[
    {
        "image_filename": "designv11_13_0000642_tits.2014.2340020-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000642_tits.2014.2340020-Figure3-1.png",
        "caption": "Fig. 3. Rotation center for the case where the vehicle rotates on a spot.",
        "texts": [
            " We also show that our proposed algorithm can generate a collision-free path, if a path is possible, with the smallest rotation and utilizes the characteristics of LPA\u2217 to quickly replan a collision-free path (see Section VI). In order to check whether any interference occurs against obstacles when a differential wheeled vehicle with differential shapes rotates on a spot, it is important to obtain the rotation center of the vehicle. In [26], the kinematics of the vehicle is presented by using the instantaneous center of rotation. However, this paper focuses on the rotation center (xc and yc in Fig. 3) for the case in which the vehicle rotates on a spot. The velocity of each wheel can be depicted, as shown in Fig. 3. The resultant velocities in the longitudinal direction (x-axis) for the left and right wheels, i.e.,VL and VR, respectively, and the lateral direction (y-axis) for the front and rear wheels, i.e., VF and VB , respectively, can be expressed as VL = v1x + v2x VR = v3x + v4x VF = v2y + v3y VB = v1y + v4y (2) where v1x, v2x, v3x, and v4x are the velocities in the longitudinal direction for each wheel, and v1y , v2y , v3y , and v4y are the velocities in the lateral direction for each wheel. For the condition in which the vehicle rotates on a spot, the resultant velocity in the longitudinal direction, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003765_iros.2018.8594316-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003765_iros.2018.8594316-Figure6-1.png",
        "caption": "Fig. 6: Detail of the thumb.",
        "texts": [
            " The way of fabrication of fingers is shown in Fig. 5. The thickness of PET plates is naturally adjusted by heating in a oven. The length and tension of ligaments is adjusted by twisting wire. As mentioned in previous section, the CM joint has two remarkable characteristics. \u2022 It has two DOFs: opposition and abduction-adduction. \u2022 The DOF of opposition has wide range of motion. To realize these characteristics in the limited space, this joint consists of three machined springs as shown in the top of Fig. 6. The CM joint is in the position of a red circle, and it is connected to the palm in the position of blue circles. The schematic diagram of spring arrangement is shown in the bottom of Fig. 6. The two of three springs are placed in parallel and this pair and the third spring are connected in series, turned back in the connected point. The pair of parallel springs are bended when oppositing and the third spring are bended when both oppositing and adducting. In the other words, when adducting only one spring is bended as the flexion of other general joints, but when oppositing two serial pairs of springs are bended. In this way the joint with wide range of motion is realized without little affection to other joints"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003714_s12008-018-0520-6-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003714_s12008-018-0520-6-Figure10-1.png",
        "caption": "Fig. 10 Effect of varying the laser beam speed on the temperature distribution of the part",
        "texts": [
            " The mid-point of layer 4 has experienced a similar thermal cycle, where the maximum temperature has a value of 2113 \u00b0C. The temperature peaks at points B and C indicate that the melting point of the material has been reached. An interesting observation is noticed in Fig. 8 that after the deposition of the fourth layer, the mid-point temperature of the 2nd layer reached 1650 \u00b0C. This aging and tempering effects may affect the mechanical properties of the part including the residual stresses and its strength. Figure 10 demonstrates the effect of varying the speed of the laser beam on the temperature distribution of the part after completing building the last layer (4th layer). It can be seen from Fig. 10 that as the velocity of the laser beam increases, the temperature of the part decreases comparedwith the case of lower velocity. This is can be attributed to less time of exposure to the laser beam. Figure 11 shows the temperature variation along the height of the part for various laser travel velocities. The plotted data was taken at the time instant after the 4th layer was deposited. The corresponding maximum temperature were about 2105 \u00b0C, 2010 \u00b0C, 1700 \u00b0C and 1510 \u00b0C for tested scan speed of \u03c0 mm/s, 2\u03c0 mm/s, 4\u03c0 mm/s and 6\u03c0 mm/s, respectively. Figure 10 clearly shows that the temperature along the height of the part is the lowest for the higher velocity (6\u03c0 mm/s) and this is due to the shorter interaction time between the laser beam and the material layer during the scanning process. However, the lowest travel velocity exhibits themaximum temperature. This paper presented a finite element model of the thermal aspects of additive manufacturing process (AM). Simulation results were validated against the experimental results found in the literature [4, 5] and showed good agreement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002375_ecc.2015.7330782-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002375_ecc.2015.7330782-Figure1-1.png",
        "caption": "Fig. 1. Relative kinematics of UAV-target motion.",
        "texts": [],
        "surrounding_texts": [
            "Cooperative target tracking of a moving ground vehicle using multiple UAVs was achieved and presented in our early work [5], where the cooperative target tracking control law was designed to include two components: \u03c8\u0307i(t) = uit(t) + uic(t). (3) The term uit(t) denotes the control effort to regulate the 2D horizontal range between each UAV and the target to the prescribed distance \u03c1d(t). The term uic(t) spreads all agents evenly on a circumference, the center of which tracks the moving target. Design of uit(t) will be given in Sec. III where the 2D horizontal range can be a time-varying reference instead of a constant. In the design of uic(t), we apply the gradientbased coordination algorithm in [5], [9] that can be achieved by using only a simple visual measurement, i.e., the bearing angles \u03b2ij(t) between the agents. The bearing angle \u03b2ij(t) is the angle between the ith UAV\u2019s velocity vector and the vector pointing from the position of the ith UAV to the position of the jth UAV. Please see [5], [9] for a graphic illustration of the bearing angle. Regarding communication topologies, three communication topologies are investigated. The coordination control laws are summarized below [5]: 1) (Undirected) All-to-All Communication: Each agent communicates with all other agents. The coordination control law takes the following form: uic(t) = \u2212\u03ba n\u2211 j=1,j 6=i cos(\u03b2ij(t)), \u03ba > 0. (4) 2) (Directed) Ring Topology: The agent i connects to agents i + 1 (modulo n) and i \u2212 1 (modulo n). The coordination control law is: uic(t) = \u2212\u03ba[cos\u03b2i(i+1)(t) + cos\u03b2i(i\u22121)(t)], \u03ba > 0. (5) 3) Cyclic Pursuit Strategy: Each agent only pursues its \u201cleading\u201d neighbor, i.e., the agent i + 1 (modulo n): uic(t) = \u2212\u03ba ( cos\u03b2i(i+1)(t)\u2212 cos \u03c0 n ) , \u03ba > 0. (6)"
        ]
    },
    {
        "image_filename": "designv11_13_0003620_j.matpr.2018.06.241-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003620_j.matpr.2018.06.241-Figure5-1.png",
        "caption": "Fig. 5 Bicycle subjected to condition",
        "texts": [
            " (ii)Steady State pedalling- the cyclist is seated on the bicycle and applying a force of 200N due to leg dynamics. The load is assumed to be concentrated at the bearing as shown in above Fig 3 (i -ii). (iii)Vertical Impact- vertical impact loads are represented by multiplying the cyclist\u2019s weight by some amount of G factor. In this case a factor of 2G is taken taking the load to 1400N which is the necessary case when an object falls from an infinitesimal height onto a rigid surface as shown in below Fig 4 and Fig 5. (iv)The case that is presented now is from [2]. Here the loads are simulated for the load bumps occurring at the front wheel. A resultant load of 2700N is transmitted at the rigid links which are then connected to the axle and then to the frame via the fork. 18924 Devaiah B.B et al./ Materials Today: Proceedings 5 (2018) 18920\u201318926 The frame is divided into 169770 triangular elements. The stress results for the different loading conditions are shown below for (i) (ii) and (iii) cases. Devaiah B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002173_detc2014-35099-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002173_detc2014-35099-Figure7-1.png",
        "caption": "Figure 7 Adaptive slicing due to a rise in surface complexity angle",
        "texts": [
            "   1 12 VE t A Ai i   (5) In developed procedure user can specify a maximum allowable volumetric dissimilarity (much like a dimensional tolerance) for the object. Then the maximum allowable volumetric dissimilarity is computed based on the cross sectional area information and the variable layer thickness and it is used for the adaptive slicing procedure to improve the surface quality of LM parts. The volumetric error control algorithm is implemented in MATLAB-12 using the GUI for visualization. Initially, critical parameters (refer Fig. 7) are defined as follows: VE = Volumetric error bound Zmin = Minimum part height Zmax = Maximum part height VEcurrent = Volumetric error due to current slicing lt = Layer thickness ltmin = Minimum layer thickness ltmax = Maximum layer thickness ds = Step (= 0.001) Typical flowchart describing the implementation procedure has been given in Fig. 8. In the present program, volumetric error which needs to be modified on the basis of user specified volumetric error bound, need some basic inputs based on 5 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings",
            " These inputs are further used in the proposed algorithm to generate variable layer thickness based on user specified volumetric error bound. In the first step, the system finds out the maximum height (Zmax) and minimum height (Zmin) of the tessellated model at given orientation. After this, algorithm starts slicing process at the minimum Z-height of the given part and uses coordinate values of the present slice plane to extract the sectional area. In second step, slice plane incremented in the direction of z-height by maximum slice thickness value ( refer Fig. 7) and determines the cross sectional area of this slice. Corresponding volumetric loss has been calculated between these two successive layers. If the value of volumetric loss at the current slice thickness is more than the user defined acceptable volumetric loss then algorithm temporarily stores these values and decreases present slice thickness by ds and repeats the process until volumetric error is within the user defined volumetric error bound or slice thickness becomes equal to minimum slice thickness ( refer Fig. 7). If slice thickness reaches to minimum slice thickness then corresponding volumetric error values are temporarirly stored by the program and from temporary stored data, program further checks all the slice thickness and searches minimum volumetric error (VEmin) among all the corresponding sliced layers. At this stage, i.e. at current Z- height new slice thickness, cross sectional area information, volumetric error and corresponding temporary minimum Z value (Zmin) are saved by the program. In the third step, the algorithm checks the difference between the current slice thickness and maximum Z-height of the given CAD model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001146_0954406217722380-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001146_0954406217722380-Figure3-1.png",
        "caption": "Figure 3. Rotor system with flexible supports: (a) sSketch of rotor system; (b) cross-view of the support assembly.",
        "texts": [
            " Dynamic models of rotor system and bearing cage The shaft vibrations excited by rotor unbalance can bring out the time-varying reaction forces on the supports, which can be eventually transmitted from the inner race to other bearing components including the cage. In order to predict the cage motions due to rotor unbalance theoretically, the dynamic models of the rotor system and the bearing cage are presented. Dynamic equations of the rotor system The rotor system for the scaled test rig has two support assemblies including angular contact ball bearing, bearing house, squirrel cage, as shown in Figure 3. The stiffness of support assembly mainly depends on the squirrel cage, because it is much more compliant than other parts. And the stiffness of supports can be obtained by referring to Wang et al.33 The dynamic model for the rotor system is illustrated in Figure 4(a). Its two flexible supports are located at point B1 and B2 on the short rigid shaft and the disc is located at point C. The horizontal and vertical stiffness of point B1 and B2 are k1y, k1z, k2y, k2z. The distances among them are represented as l, a, b, also shown in Figure 4(a)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure2.2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure2.2-1.png",
        "caption": "Fig. 2.2 Conditional distribution of the fluxes of mutual and full self-induction of the induction machine windings (a) fluxes of mutual and full self-induction of the stator winding; (b) fluxes of mutual and full self-induction of the rotor winding; (c) fluxes of mutual and full self-induction of the stator and rotor windings",
        "texts": [
            " 15 The emfs E1 and E2 represent the total fluxes\u03a61 and\u03a62 taking place in an electric machine under the unilateral power supply of its windings, and the emfs E12 and E21 reflect the fluxes of mutual induction\u03a612 and\u03a621, respectively (Fig. 2.1a, b). In this work, we consider a symmetrical electricmachinewith symmetrical power supply in its windings. Therefore, for ease consideration, the area of the magnetic field distribution can be limited to the single tooth division. The conditional magnetic field images corresponding to the equations (2.3) in Fig. 2.1 can then be represented in a more schematic and visual form in Fig. 2.2. In these figures, the stator and rotor teeth regions are subdivided into two areas: areas occupied by the windings and areas where slot wedges are located. In this case, the air gap is considered to comprise the two sub-layers with identical thicknesses equal to \u03b4/2. This provision is discussed below in greater detail. On the basis of expressions (2.4), (2.5), and (2.6), the resulting emfs of the stator and rotor windings (E1\u0440 and E2\u0440) take the form E1\u0440 \u00bc E1\u00f0 \u00de \u00fe E21\u00f0 \u00de \u00bc jx1I1 \u00fe jx21I2 E2\u0440 \u00bc E12\u00f0 \u00de \u00fe E2\u00f0 \u00de \u00bc jx12I1 \u00fe jx2I2 \u00f02:7\u00de The system of equations (2.7) corresponds to the resulting fluxes \u03a61p and \u03a62p, as determined by the equations (2.3). On the basis of the system of equations (2.7), the conditional picture of the resulting flux distribution receives the form in Fig. 2.2c. According to Fig. 2.2c, the stator and rotor windings are considered to be two inductively coupled windings, and the system (2.7) in this case represents the emf equations for these two inductively coupled windings. 16 2 Fundamentals of the Field Decomposition Principle The equations (2.7) can be used to determine the emfs of the stator and rotor windings. For example, from (2.7), we can obtain the emf of the stator winding E1\u0440 \u00bc jx1 1 x12x21 x1x2 I1 x21 x2 E2\u0440 \u00f02:8\u00de In (2.8), determining the reactance values x1, x2, x12 and x21 for a real electric machine with the accuracy that is required represents a very difficult task"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003599_j.procir.2018.04.033-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003599_j.procir.2018.04.033-Figure8-1.png",
        "caption": "Fig. 8 Representation of the measurement processes: (a) massive and (b) lattice specimen",
        "texts": [
            " The massive specimen is used to measure the height variation. The lattice structure is used to observe the variation of the geometry of the cell. The authors have obtained different kinds of defects and they concentrate their study on the distortion of the first layers, see Fig. 7. The distortions of the first layers of 25 massive specimens were measured by a three-dimensional optical control machine (Vertex from Micro-vu). This equipment is able to measure the height of each point in the first layer, see Fig. 8a. The difference between the nominal and the measured height is computed. The lattice structure defect is observed using a different measurement protocol, see Fig. 8b. A gauge tool was created to detect a \u00b110 % change in the shape of every cell of the part. The different height variations of massive test specimens are presented in Fig. 9. An example of the height defect of one specimen is shown in Fig. 9a: the height defect is presented in function of the angular position of measured point. The defects cartography of 25 specimens is shown in Fig. 9b. It could be observed that the defect value is more important when the measured point on a specimen is close to the borders of the platform of the machine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003650_s12206-018-0907-0-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003650_s12206-018-0907-0-Figure3-1.png",
        "caption": "Fig. 3. Displacements and rotations in the joined conical-cylindrical shells.",
        "texts": [
            " (19) are defined as { } { } { } { } 1 2 1 2 1 2 1 2 1 2 1 2 1 2 1 2 1 2 1 2 . Tcy cy cy cy cy cy cy cy cy cy cy cy cy cy cy P P P P P Tcy cy cy cy cy cy cy cy cy cy P P P P P P P P P P a a a b b b c c c d d d e e e a a b b c c d d e e g d + + + + + + + + + + = \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 = (20) Matrices [C1(n)], [C2(n)], [D1(n)] and [D2(n)] correspond to the circumferential wave number n. 2.3 Joined conical-cylindrical shell A joined shell structure is constructed by assembling a conical shell and a cylindrical shell together as shown in Fig. 3. Subdomains 1W and 2W represent the conical shell and the cylindrical shell, respectively. Eqs. (11) and (20) are applied to the conical and the cylindrical section separately, however, the equations of motion are written together as ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) 1 2 1 2 1 22 1 2 0 0 0 0 0 0 . 0 0 n n n n n n n n b c g d b c w g d \u00ec \u00fc \u00ec \u00fc\u00e9 \u00f9 \u00e9 \u00f9\u00ef \u00ef \u00ef \u00ef+\u00ea \u00fa \u00ea \u00fa\u00ed \u00fd \u00ed \u00fd \u00ea \u00fa \u00ea \u00fa\u00ef \u00ef \u00ef \u00ef\u00eb \u00fb \u00eb \u00fb\u00ee \u00fe \u00ee \u00fe \u00e6 \u00f6\u00ec \u00fc \u00ec \u00fc\u00e9 \u00f9 \u00e9 \u00f9\u00ef \u00ef \u00ef \u00ef\u00e7 \u00f7= +\u00ea \u00fa \u00ea \u00fa\u00ed \u00fd \u00ed \u00fd\u00e7 \u00f7\u00ea \u00fa \u00ea \u00fa\u00ef \u00ef \u00ef \u00ef\u00eb \u00fb \u00eb \u00fb\u00ee \u00fe \u00ee \u00fe\u00e8 \u00f8 A A C C B B D D (21) The total number of equations in Eq",
            " 0 0 n n n n n n n n n n n n b g bq w g - - \u00e6 \u00f6\u00ec \u00fc\u00e9 \u00f9 \u00e9 \u00f9 \u00ef \u00ef\u00e9 \u00f9 \u00e9 \u00f9\u00e7 \u00f7-\u00ea \u00fa \u00ea \u00fa \u00ed \u00fd\u00eb \u00fb \u00eb \u00fb\u00e7 \u00f7\u00ea \u00fa \u00ea \u00fa \u00ef \u00ef\u00eb \u00fb \u00eb \u00fb \u00ee \u00fe\u00e8 \u00f8 \u00e6 \u00f6\u00ec \u00fc\u00e9 \u00f9 \u00e9 \u00f9 \u00ef \u00ef\u00e9 \u00f9 \u00e9 \u00f9\u00e7 \u00f7= -\u00ea \u00fa \u00ea \u00fa \u00ed \u00fd\u00eb \u00fb \u00eb \u00fb\u00e7 \u00f7\u00ea \u00fa \u00ea \u00fa \u00ef \u00ef\u00eb \u00fb \u00eb \u00fb \u00ee \u00fe\u00e8 \u00f8 A A Z Z X X B B Z Z Y Y (41) The solution of Eq. (41) yields the estimate for the natural frequencies and corresponding mode shapes of the joined conical-cylindrical shells. 3. Numerical examples 3.1 Joined conical-cylindrical shell structure A joined shell structure is constructed by joining a conical shell and a cylindrical shell together as shown in Fig. 3. The geometric characteristics of the joined conical-cylindrical shell structure are A/R = 0.8452, B/R = 1.1548, L/R = 1, h/R = 0.01 and / 6.a p= Poisson\u2019s ratio and the shear correction factor are n = 0.3 and k = 5/6. To investigate the accuracy and the convergence property of the present method the natural frequencies of the joined conical-cylindrical shell with clampedclamped boundary conditions are computed for different K and P and the results are given in the dimensionless form ( )21R EW w r n= - (42) in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002322_j.triboint.2015.10.007-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002322_j.triboint.2015.10.007-Figure12-1.png",
        "caption": "Fig. 12. Installation of the mechanical seal in the holder.",
        "texts": [],
        "surrounding_texts": [
            "According to the experimental results, the three examined types of shaft seals appeared to be in the severe mixed lubrication regime. In the water-sealing system, the frictional torque was not affected by water pressure variations in the range 0.3\u20130.8 MPa (Fig. 15), and no obvious change was observed in the amplitude variability and the value of the frictional torque after changing the water pressure (Fig. 16). The friction independent of contact pressure has been often observed in the boundary lubrication mode. In the oil seal and mechanical seal, wear and damage were detected on the dynamic seal face (Figs. 17 and 18), and the results were acquired using a low shaft rotational speed. The low speed is considered to inhibit lubricating film formation between the dynamic seal faces. The water-sealing system has two distinct characteristics: the use of two hydrated seal lips and the application of a nonNewtonian fluid between them. PVF has numerous hydroxyl groups and a continuous porous structure, and thus the PVF seal lips were expected to produce hydration lubrication between the dynamic seal faces. If a seal lip made of hydrophobic rubber is adopted, as in an oil seal, no formation of a lubricating film generated by the water is expected. Therefore, high adhesion between the seal surfaces of the oil seal was demonstrated in this study (Fig. 17). Although hydration lubrication was expected when using the PVF seal lip, the viscosity of the water in the PVF and between the dynamic seal faces was too low to maintain a lubrication film between the dynamic seal faces. This is the first reason why an aqueous PEG solution was adopted as the lubrication liquid in this study. As shown in Fig. 6, the PEG solution has non-Newtonian characteristics, meaning the viscosity decreases with increasing shear rate. The PEG solution present between the two seal lips, where its shear rate appears to be low, may act as a high-viscosity material. This characteristic affects the prevention of water ingress from the water phase to the air phase. The PEG solution between the dynamic seal faces, where the shear rate of the PEG appears to be high, may act as a low-viscosity lubricant because this lubricating film is thin and the relative motion between the faces is large. This leads to the increased lubricant entrainment between the dynamic seal faces and the reduction of frictional torque originating from the shear stress of the lubricant. The lubricant entrainment between the dynamic seal faces may behave well with an associated characteristic of the hydrated seal lips, because the results shown in Fig. 14 indicate that no pressurization of the lubricant is required between the two seal lips. As shown in Fig. 15, the water-sealing system exhibited lower frictional torque than the oil and mechanical seals and its friction was unaffected by the applied water pressure. The observed independence from the pressure may be owing to the excellent compatibility between the hydrated material and the water-based lubricant and also the flexibility of the configuration at the dynamic seal faces. The geometry of the PVF seal lip near a rotating shaft was a simple rectangular shape and its elastic modulus was approximately 9.0\u201310.0 MPa. When misalignment of the seal lip and the rotating shaft occurred because of the water pressure, adequate deformation of the seal lip was expected to establish the appropriate sealing and lubricating conditions at the dynamic seal faces. The complicated shape of the seal lip of the oil seal is the result of a considerable amount of research aimed at improving its performance; however, the misalignment between the seal lip and the rotating shaft due to the deformation of the lip by the water pressure may increase the contact pressure between the seal lip and the rotating shaft, which leads to larger amplitude variability of the frictional torque (Fig. 16). This may cause high frictional torque and excessive abrasion of the lip, ultimately leading to water leakage (Figs. 15 and 17). In the mechanical seal, the dynamic seal faces consist of two rings. The rings are pressed against each other by a spring to improve the conformity between the sliding surfaces. Because of this structure, the contact pressure between the two bearing surfaces was increased by increasing the water pressure (Fig. 16), which resulted in increased frictional torque (Fig. 15). This high contact pressure sometimes damaged the frictional surface, as shown in Fig. 18. Although it was confirmed that the water-sealing system can exhibit low frictional torque and low water leakage for low shaft rotational speed, the limitation of shaft speed and its diameter seems to be found. Limitation of temperature seems also to be found, because the heatproof temperature of PVF is lower than 80 \u00b0C. The influence of added substances in the water should also be investigated because these may induce chemical reactions with PVA or PEG. In addition to durability testing, the above issues are considered for future investigation. The cost of manufacturing for the water-sealing system is estimated to be lower than that for conventional shaft seals because the raw materials, such as PVA and PEG, are not expensive and the housing for the installation of the hydrated seal lips is not very complicated."
        ]
    },
    {
        "image_filename": "designv11_13_0000564_j.cja.2013.12.010-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000564_j.cja.2013.12.010-Figure1-1.png",
        "caption": "Fig. 1 Schematic of the spacecraft with a rotating flexible solar array.",
        "texts": [
            ", orbital dynamics will not be considered in attitude dynamics); (ii) the center of mass of the spacecraft will not change in spite of the rotations and vibrations of the solar array; and (iii) variables of the attitude motion of the spacecraft, and those of the rotations and vibrations of the solar array are all first-order infinitesimal variables such that their products are high-order infinitesimal terms and can be neglected in the mathematical model. Let x = [xx xy xz] T 2 R3, X 2 R, and g = [g1 g2 . . . gN] T 2 RN denote respectively the angular velocities of the spacecraft, the angular velocity of the solar array, and modal coordinates with truncation number N, one of the well-known formulations of the attitude dynamics of the spacecraft system, as shown in Fig. 1, is then presented as follows10,11 J _x\u00fe Jbsl _X\u00feHbs\u20acg \u00bc T\u00feHOT1\u00f0x;X; _g; g\u00de lTJTbs _x\u00fe lTIsl _X\u00fe lTHs\u20acg \u00bc s\u00feHOT2\u00f0x;X; _g; g\u00de HT bs _x\u00feHT s l _X\u00fe \u20acg \u00bc D _g Kg\u00feHOT3\u00f0x;X; _g; g\u00de 8><>: \u00f02\u00de where HOT1\u00f0x;X; _g; g\u00de, HOT2\u00f0x;X; _g; g\u00de and HOT3\u00f0x;X; _g; g\u00de are the high-order infinitesimal terms that do not contain any acceleration variables \u00f0 _x; _X, or \u20acg\u00de. l = [0 1 0]T is the rotational direction of the solar array, which means the solar array is rotating along the pitch axis of the spacecraft. The N \u00b7 N diagonal matrices D and K are the orthonormal modal damping and stiffness of the flexible appendage, and both are strictly positive definite",
            " The inertia matrices of the central body and the solar array are given by Jb \u00bc 1028:8716 10:9150 23:6426 10:9150 1026:2444 17:8563 23:6426 17:8563 1117:2465 264 375 kg m2 and I s \u00bc 1104:6067 9:3024 0:0576 9:3024 28:7784 2:2498 0:0576 2:2498 1076:3055 264 375 kg m2 The flexible coupling matrices of the solar array are H s \u00bc 0:001312 33:233892 0:002936 0:326503 0:006997 0:009509 0:329167 0:067510 4:868742 0:116229 31:935654 0:001202 5:618361 0:128195 3:833407 264 375 ffiffiffiffiffi kg p m and Ps \u00bc 6:362845 0:000428 3:745014 0:139602 3:858466 0:060110 0:061114 0:346686 0:007505 0:220996 0:000268 7:519367 0:001814 0:196472 0:000649 264 375 ffiffiffiffiffi kg p Both of them consist of the influence of five flexible modes, namely, the first, third, and fifth modes that are mainly out-of-plane bending, the second mode that is mainly in-plane bending, and the fourth mode that is mainly torsion. Natural frequencies of these modes are K = 2p diag(0.172617, 0.692549,1.072639,1.832541,2.507761) rad/s, and the stiffness matrix is K= K2. Supposing damping ratios are n = diag(0.004,0.005,0.0064,0.008,0.0085), the damping matrix can be generated by D= 2nK. The two position vectors about the solar array are b= [ 0.02764 0.78871 0.00844]T m, and rs = [ 0.02045 2.91590 0.02455]T m, where rs is the position vector from the hinge point to the mass center of the solar array w.r.t. Fs (see Fig. 1). With the mass ms = 85.74 kg, the first moment of inertia of the solar array can be calculated by cs = msrs. Using ms, cs, and Is, one can investigate the inertial completeness of the flexible coupling matrices by modal identities (12). For simplicity, the initial attitude error q(0) and rotation angle error Da(0) are presumed to be [1 0 0 0]T and zero respectively. Suppose that the orbital angular velocity is Xo = 0.060 optimal control system. ( )/s, then the desired angular velocities are given by xd = [0 Xo 0]T, and Xd = Xo"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000616_aim.2012.6265962-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000616_aim.2012.6265962-Figure2-1.png",
        "caption": "Fig. 2. Gaits including floating-foot",
        "texts": [
            " (9) Finally, we get motion equation with one leg standing as: M(q)q\u0308 + h(q, q\u0307) + g(q) + Dq\u0307 = \u03c4 , (10) where, M(q) is inertia matrix, h(q, q\u0307) and g(q) are vectors which indicate Coriolis force, centrifugal force and gravity, D = diag[d1, d2, \u00b7 \u00b7 \u00b7 , d18] is matrix which means coefficients of joints\u2019 viscous friction and \u03c4 is input torque. If supporting-foot is surface-contacting and assumed to be without slipping, joint angle can be thought as q = [q2, q3, \u00b7 \u00b7 \u00b7 , q18]T . This walking pattern is depicted in Fig. 2 (a). When heel of supporting-foot should detach from the ground before floating-foot contacts to the ground as shown in Fig. 2 (b), the state variable for the foot\u2019s angle q1 be added to q, thus q = [q1, q2, \u00b7 \u00b7 \u00b7 , q18]T . Giving floating-foot contacts with a ground, contactingfoot like Fig. 3 appears with contacting-foot\u2019s position zh or angle qe to the ground being constrained. When constraints of foot\u2019s position and also foot\u2019s rotation are defined as C1 and C2 respectively, these constraints are represented by Eq. (11), where r(q) means the contacting-foot\u2019s heel or toe position in \u03a3W . C(r(q)) = [ C1(r(q)) C2(r(q)) ] = 0 (11) Then, robot\u2019s equation of motion with external force fn, friction force ft and external torque \u03c4n corresponding to C1 and C2 can be derived as: M(q)q\u0308 + h(q, q\u0307) + g(q) + Dq\u0307 = \u03c4 + jT c fn \u2212 jT t ft + jT r \u03c4n, (12) where jc, jt and jr are defined as: jT c = \u201e \u2202C1 \u2202qT \u00abT\u201e 1/ \u201a\u201a\u201a\u201a \u2202C1 \u2202rT \u201a\u201a\u201a\u201a \u00ab , jT t = \u201e \u2202r \u2202qT \u00abT r\u0307 \u2016r\u0307\u2016 , (13) jT r = \u201e \u2202C2 \u2202qT \u00abT\u201e 1/ \u201a\u201a\u201a\u201a \u2202C2 \u2202qT \u201a\u201a\u201a\u201a \u00ab ",
            " Here, fn and \u03c4n are decided dependently to make the q\u0308 in Eq. (12) and Eq. (15) be identical. 2 4 M(q) \u2212(jT c \u2212 jT t K) \u2212jT r \u2202C1/\u2202qT 0 0 \u2202C2/\u2202qT 0 0 3 5 2 4 q\u0308 fn \u03c4n 3 5 = 2 6664 fi \u2212 h(q, q\u0307) \u2212 g(q) \u2212Dq\u0307 \u2212q\u0307T j \u2202 \u2202q \u201e \u2202C1 \u2202qT \u00abff q\u0307 \u2212q\u0307T j \u2202 \u2202q \u201e \u2202C2 \u2202qT \u00abff q\u0307 3 7775 (16) Here, since motion of the foot is constrained only vertical direction, walking direction has a degree of motion. That is, contacting-foot may slip forward or backward depending on the foot\u2019s velocity in traveling direction. As shown in Fig. 2, we distinguish contacting patterns by changing the dimension of state variables. That is, although we do not address the situation that supporting-foot slips or both feet are in the air, Eq. (16) can also represent these dynamics: adding position variable of walking direction y0 to q in Eq. (16) when supporting-foot begins slipping; and jumping motion by adding further variable of upright direction z0 to q when jumping represented by Fig. 4. Table II indicates all possible walking gaits regarding contacting situations\u2014surface-contacting (S), point-contacting (P) and Floating (F)\u2014of supporting-foot (S"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001204_icma.2017.8015804-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001204_icma.2017.8015804-Figure7-1.png",
        "caption": "Figure 7 Emitter Circuit of the System.",
        "texts": [
            " To construct the architecture shown in figure 6 the HMI receive the image from the camera, the signals of the car and the signals from the driver, the signals from the driver are transformed into digital and analogic signals that are transmitted, between the PC and the car there is an emitter circuit whose manage the analogic controls data, reading its voltage levels and sending the data to the vehicle and PC\u2019s serial port. For communicating with the PC, a PL2303 TTL to USB converter was used, TTL pins connect directly to the module in the car. Wireless communication is managed by a HC-05 Bluetooth module. It connects directly to the receptor Bluetooth module located on the vehicle. Other pins are connected to the external devices to be read and two motors to control vibration. Figure 7 shows the schematic diagram for the connections. The receptor circuit located on the car has an HC-06 Bluetooth module for receiving data, which is directly sent to the Microcontroller for its decoding. Other pins are connected to the external actuators such motors and lights. For controlling the principal motor and its speed, pulse width modulation (PWM) signal was used depending on the data received. Battery is connected to a 5v regulator for feeding the circuit, and is connected to a 12 volts motor driver, for driving the electric DC motor, as we can see in the Figure 8"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000647_iciea.2015.7334244-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000647_iciea.2015.7334244-Figure11-1.png",
        "caption": "Fig. 11 Simulation system for manipulator",
        "texts": [
            " Therefore, there's a collision detection list for every geometry, and we only need to process geometry in the list when traversing collision detection. The model system of a manipulator is shown in Fig. 10. V. SIMULATION AND EFFICIENCY OF THE ALGORITHM As mentioned in section II, we mainly focused on manipulators of cylinder and cuboid shape and the algorithm is in high precision under this condition. In order to verify the validity of the algorithm, we developed a simulation system based on OSG for manipulator, as shown in Fig. 11. The D-H parameters of the manipulator are given in Table 3. The 8 links from the base to the end is defined L1 to L8, and L0 represents the base. Slide guide S and Flatbed F are seen as environment. Collision detection works rapidly by adopting the algorithm in this paper. To verify the efficiency of the algorithm, we set many basic geometry of different postures for collision detection. Each pair of geometries is tested 1 million times so as to compute how long time a single collision detection spends"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001865_b978-0-12-374920-8.00808-0-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001865_b978-0-12-374920-8.00808-0-Figure14-1.png",
        "caption": "Figure 14 Three different models for the nature of the ICT state in peridinin and other carbonyl-containing carotenoids dissolved in polar solvents: (left) S1 and ICT strongly quantum mechanically mixed (S1/ ICT), (middle) ICT as a separate electronic state from S1 (S1 \u00fe ICT), and (right) S1 and the ICT state one and the same (S1(ICT)). Solid lines represent radiative transitions. The dashed line represents",
        "texts": [
            " This behavior is highly anomalous for carotenoids, which generally show very little dependence of the spectral properties and life- times of the S1 excited singlet state on the solvent environment. The presence of an ICT state in the excited state manifold was proposed to explain these observations,92,93 and this idea has been supported by both theoretical computations94 and experiments on numerous other carbonylcontaining carotenoids and polyenals.95\u2013104 However, whether the ICT state is quantum mechanically mixed with S1 (S1/ICT in Figure 14),97,105 a separate electronic state from S1 (S1\u00fe ICT in Figure 14),92,94,106,107 or simply the S1 state with a large intrinsic dipole moment brought about by mixing with the S2 state (S1(ICT) in Figure 14)108 remains uncertain. In any case, these data and results from Stark spectroscopy109 suggest that alterations in the polarity of the protein environment in the vicinity of the bound peridinin molecules may provide a mechanism through which light harvesting may be regulated in antenna complexes containing carbonyl carotenoids.110 An important aspect of the PCP complex is that it is amenable to both site-directed mutagenesis and, as mentioned previously, reconstitution with different Chls.91 These transformations alter the amino acid composition and posi- tions of the donor and acceptor absorption bands, allowing a systematic exploration of the controlling features for energy transfer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003620_j.matpr.2018.06.241-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003620_j.matpr.2018.06.241-Figure2-1.png",
        "caption": "Fig. 2 Bicycle model with rigid links",
        "texts": [
            " These rigid bodies are force transmitting members and are fixed to front fork and rear axle. The seat post and the steering tube were also simulated as rigid links. A similar study was also performed by Pazare et al [3] , where the frame was considered as a truss to find the stresses theoretically in each member and then to compare it with FE results using commercial package Ansys APDL. The tubes were assumed as line i.e. beams elements and subjected to loads as mentioned in previous literatures as shown in below Fig 1 and Fig 2. The approach involves in validating two existing research papers [1] and [2]. The former approach involves just the frame where the loads are applied directly to the frame members. The loads being taken from the references [4] and [5]. The different tubes are named and their respective diameters and length are also given. A uniform thickness of 1mm is given to all the tubes and radius of 4mm is maintained at all the junction of the tubes. Whereas, in the latter approach the loads are applied at the rigid links to simulate real world conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000652_lmag.2012.2214027-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000652_lmag.2012.2214027-Figure3-1.png",
        "caption": "Fig. 3. (a) Experimental apparatus used for measuring the force between a permanent magnet and an identical permanent magnet or an iron plate, both attached to (b) nonmagnetic holders.",
        "texts": [
            " Note that, since a displacement by dz of the magnet with respect to the plate implies a displacement of 2dz between the magnet and its image, the force between the magnet and its image is the same as the force between two magnets, even if their interaction energies differ by a factor of two. In order to validate the theoretical framework, we measured the force as a function of the distance 1) between three NdFeB axially magnetized cylindrical permanent magnets and a soft iron plate with dimensions 25 mm \u00d7 25 mm \u00d7 10 mm, and 2) between the same NdFeB magnets and their identical copies. The force was measured with a tensile testing machine, fully computerized and equipped with a 100-N load cell (see Fig. 3). The three permanent magnets considered were cylinders with 1) R = 2.5 mm, d = 0.75 mm (\u03c4 = 0.3), \u03bc0 M0 = 1.07 T, 2) R = 2.5 mm, d = 2.5 mm (\u03c4 = 1), \u03bc0 M0 = 1.12 T and 3) R = 2 mm, d = 5 mm (\u03c4 = 2.5), \u03bc0 M0 = 1.22 T. The saturation magnetization of the plate was measured with Vibrating Sample Magnetometry to be 1.66 MA/m, with a relative permeability around 1400. The magnets\u2019 saturation was measured by positioning a Hall probe at the center of their pole and correcting the readout value BH P by the geometrical factor representing the ratio between the field at the center of the exit pole and the axial field of an infinite cylindrical magnet: \u03bc0 M0 = BH P (4 + \u03c4\u22122)1/2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000912_tfuzz.2015.2466111-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000912_tfuzz.2015.2466111-Figure3-1.png",
        "caption": "Fig. 3. Time-varying nonlinearities f(e, t) belonging to the sectors: a) [0, k], b) [\u03b2, \u03b2 + k]",
        "texts": [
            " The circle criterion assumes that the control system consists of a linear dynamic element and a nonlinear static element which function is a sector bounded nonlinearity. It will be further shown that with the proper selection of consequents, the fuzzy controller function may fulfill the requirement of the sector condition. 1) Sector nonlinearity: We assume that the nonlinear timevarying function u(t) = f [e(t), t] of the adaptive fuzzy controller satisfies the following sector condition [105]: 0 \u2264 f(e, t) e \u2264 k, f(0, t) = 0, (3) which means that it lies between the lines u = 0 and u = ke (Fig. 3a). It can thus be said that the controller function is the sector nonlinearity or that it lies in the sector [0, k]. If the condition (3) is true for any e \u2208 (\u2212\u221e,\u221e), the system is said to be absolutely stable. The case is also considered where the function f(e, t) lies in the sector [\u03b2, \u03b2 + k], where \u03b2, k \u2265 0. This means that the condition \u03b2 \u2264 f(e, t) e \u2264 \u03b2 + k, f(0, t) = 0 (4) is satisfied and the function lies between the lines u = \u03b2e and u = (\u03b2 + k)e, which is shown in Fig. 3b. In this case the following substitution is used (see, for example [106]): u1(t) = f1(e, t) = f(e, t)\u2212 \u03b2e(t), (5) which is illustrated in Fig. 4. After applying the transformation (5) the function u1(t) = f1[e(t), t] lies in the sector [0, k], so the following condition is satisfied: 0 \u2264 f1(e, t) e \u2264 k, f1(0, t) = 0. (6) It is worth noting that the transformation (5) can be applied to the systems where the plant is unstable. In this case, the linear closed-loop system is stable if a value of \u03b2 exists, for which the transformed transfer function G1(s) = G(s) 1 + \u03b2G(s) (7) is stable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001583_access.2018.2868497-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001583_access.2018.2868497-Figure1-1.png",
        "caption": "FIGURE 1 The coordinate systems of the AACMM",
        "texts": [
            " The basic principles are given as follows: (1) One determines the z axis of each coordinate system and the iz axis goes along the axial direction of the joint 1i + . (2) One determines the origin O of each coordinate system and selects the origin iO on the common normal of 1iz \u2212 axis and iz axis. (3) One determines the x axis of the coordinate system; ix axis goes along the common normal between iz axis and 1iz \u2212 axis and points to the leaving direction of 1iz \u2212 . (4) One determines the y axis of the coordinate system and sets i i iy x z= + . According to the above steps, the establishment of a measuring machine coordinate system can be shown in Figure.1. Then, we can obtain the kinematic parameters of AACMM. That is, the length of linkages, the length of joints, the twist angle of linkages and the angle of joints. There are six parameters contained in each group. The kinematic parameters of AACMM are determined as follows: (1) Link length, 1ia \u2212 is the distance from 1iz \u2212 to iz along 1ix \u2212 . (2) Twist angle, 1i\u03b1 \u2212 is the angle between 1iz \u2212 to iz along 1ix \u2212 . (3) Offset length, id is the distance from 1ix \u2212 to ix along 1ix \u2212 . (4) Joint angle, i\u03b8 is the angle between 1ix \u2212 to ix along iz "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001536_s00170-018-2494-8-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001536_s00170-018-2494-8-Figure2-1.png",
        "caption": "Fig. 2 Schematic illustration of a GTAW additively manufactured Ti-6Al-4V thin wall structure, and b T-fillet structure produced from wrought Ti-6Al-4V plates for side milling machining trials",
        "texts": [
            " Additionally, a custom fabricated trailing shield was used to ensure that a sufficiently large shielding envelope was generated to the rear of the fusion zone to prevent post-weld atmospheric contamination of the successive weld beads. The general arrangement of GTAW equipment is illustrated in Fig. 1. Shielding gas flow rates and other common process parameters are detailed in Table 2. Side milling was carried out on the GTAW additively manufactured Ti-6Al-4V multi-layered deposited structures. A total of four build-up structures were produced using the process variables listed in Table 3. All structures consisted of a single weld pass per layer, producing a wall width of approximately 8 mm (schematic shown in Fig. 2a). For each sample an initial \u2018skimming\u2019milling pass was performed to generate a smooth consistent surface such that the thickness of the wall was machined down to ~ 7 mm. The fabricated \u2018Wall_02\u2019was used for an initial trial to validate the machining approach and the CNC G-code.1 Moreover, to provide comparison to machining of wrought Ti-6Al-4V, a sample of similar geometry to the additively manufactured build-ups was fabricated from two sections of commercially sourced wrought Ti-6Al-4V plate measuring 200 mm \u00d7 50 mm \u00d7 12 mm. These were welded together in a T-fillet configuration (as shown in Fig. 2b) using 1 mm diameter Ti-6Al-4V filler wire. A \u2018double-V\u2019 preparation was used on the vertical plate to produce fully penetrated weld joint. Prior to the machining trials, the vertical wall section of this T-fillet was milled along its entire length, reducing its thickness to 7 mm to better match the geometry and rigidity of additive build-ups. Milling trials were conducted in collaboration with Seco Tools Australia using a DMG MORI DMU 70 5-axis CNC milling machine. Samples were securely fastened to a Kistler 9257A three-component force dynamometer which was mounted to the milling machine table as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000868_physreve.90.033012-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000868_physreve.90.033012-Figure1-1.png",
        "caption": "FIG. 1. (Color online) Initial configurations of an elastic rod with three different mean curvatures. The motor point is located at the bottom of the rod and fixed in space. The motor turns counterclockwise when viewed from above at any specified frequency. {D1,D2,D3} exemplifies an orthonormal triad.",
        "texts": [
            " The Kirchhoff rod model is employed to represent a long but thin elastic body by using a space curve X(s) which expresses the centerline of the rod, and an associated orthonormal triad, {D1(s),D2(s),D3(s)} for 0 s L, where L is the length of the rod and s is a Lagrangian coordinate along the rod. The vector D3(s) is initialized as a unit tangent vector to the rod and the other two vectors are perpendicular to the tangent vector following the right-hand rule. The triad captures the amount of bending and twisting of the elastic rod. Figure 1 displays some initial configurations of the elastic rod by varying the initial mean curvature and also displays an example of the orthonormal triad. A motor point at s = 0 is embedded in the bottom of the rod and rotates by turning the triad of the motor at a prescribed angular velocity \u03c9 as follows: X(0,t) = X0, (1) D1(0,t) = (cos(\u03c9t), sin(\u03c9t),0), (2) D2(0,t) = (\u2212 sin(\u03c9t), cos(\u03c9t),0), (3) D3(0,t) = (0,0,1), (4) where |\u03c9| = 2\u03c0f/s, so f is the rotation rate in Hz. Note that \u03c9 > 0 corresponds to counterclockwise (CCW) rotation, whereas \u03c9 < 0 corresponds to clockwise (CW) rotation, assuming that the rod is viewed looking back towards s = 0 from a location at which s > 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003778_1.4042636-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003778_1.4042636-Figure14-1.png",
        "caption": "Fig. 14 Phase trajectories and Poincare maps of clearance joint: (a) phase trajectories when c 5 0.15 mm in the Y direction, (b) Poincare map when c 5 0.15 mm in the Y direction, (c) phase trajectories when c 5 0.45 mm in the Y direction, (d) Poincare map when c 5 0.45 mm in the Y direction, (e) phase trajectories when c 5 0.80 mm in the Y direction, and (f) Poincare map when c 5 0.80 mm in the Y direction",
        "texts": [
            " Figure 13 is bifurcation diagram of the velocity in the X, Y, and Z directions with the change of clearance value. From the bifurcation diagram, it can be seen that the motion state of the moving platform changes gradually from the initial periodic state to quasiperiodic state when the clearance value increases from 0.01 mm to 0.80 mm, and there is no chaos. The reason why chaos does not occur is that the clearance has a great influence on the spherical joint, but small on the moving platform. Figure 14 shows the phase trajectory and Poincare mapping of the clearance joint in the Y direction when the clearance values are 0.15 mm, 0.45 mm, and 0.80 mm. From Figs. 12(a)\u201312(d), it can be seen that the phase trajectory shows certain regularity when the clearance values are 0.15 mm and 0.45 mm, and the Poincare mapping also concentrates on the right side of the phase trajectory, which can be judged that the motion state is quasiperiodic. It can be seen from Figs. 12(e) and 12(f) that the phase trajectory lines are intertwined and crossed each other when the clearance value is 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001875_20140313-3-in-3024.00016-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001875_20140313-3-in-3024.00016-Figure1-1.png",
        "caption": "Figure 1. The Qball-X4 quadcopter UAV [4]",
        "texts": [
            " Discussion and Conclusion are presented in section VI and section VII respectively. Various practical factors can hinder the operation of quadcopters and create challenges in the control design. To achieve reliable and accurate tracking control of quadcopters, controller must be designed to compensate for model uncertainty and external disturbances. This makes it all the more challenging to implement a control design on hardware than merely designing it on paper. II. QBALL-X4 QUADCOPTER The quadcopter made by Quanser Inc., known as Qball-X4 as shown in Fig. 1, is an innovative rotary-wing aerial vehicle platform suitable for a wide variety of UAV research and development applications. The Qball-X4 is a quadcopter propelled by four motors fitted with propellers. QuaRC, Quanser\u201fs real-time control software allows researchers and developers to rapidly develop and test controllers on actual hardware through MATLAB \u00ae /SIMULINK \u00ae interface. The structure, operation and interfacing of the host machine with Qball-X4 are discussed in detail in [4]. 978-3-902823-60-1 \u00a9 2014 IFAC 192 10.3182/20140313-3-IN-3024.00016 The controllers are developed in SIMULINK \u00ae with QuaRC on the host computer, and these models are downloaded and compiled into executable codes on the target. A diagram of this configuration is shown in Fig. 2. The following hardware and software are embedded in Qball-X4:  Qball-X4: as shown in the Fig. 1.  HiQ: QuaRC aerial vehicle Data Acquisition Card (DAQ).  Gumstix: The QuaRC target computer. An embedded, Linux-based system with QuaRC runtime software installed.  Batteries: Two 3-cell, 2500 mA\u2219h Lithium-Polymer batteries.  Real-Time Control Software: The QuaRC-SIMULINK \u00ae configuration, as detailed in [4]. III. MATHEMATICAL MODEL OF THE QBALL-X4 The quadcopter considered in this work is an underactuated system with six outputs and four inputs and the states are highly coupled. Control of the quadcopter is attained by varying the speeds of each rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002012_msec2014-4029-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002012_msec2014-4029-Figure4-1.png",
        "caption": "FIG 4: GEOMETRY WITH THERMAL LOAD APPLICATION AND BOUNDARY CONDITIONS",
        "texts": [
            " The first term represents heat loss due to conduction from the surface whose unit normal is \ud45b . The second and third term refers to convection and radiation heat losses from the surface of the work-piece. For the mechanical analysis external loading is not considered and to prevent the rigid body motion the nodes on the base are fully constrained to prevent elemental motion. The application of thermal load as surface heat flux and the mechanical boundary conditions showing the nodes which are fully constrained are illustrated in Figure 4. Temperature dependent thermo-physical properties are considered for both clad and substrate. Along with thermal properties like thermal conductivity, specific heat and latent heat, mechanical properties namely co-efficient of thermal expansion, Young\u2019s modulus, Poison\u2019s ratio and yield strength are provided as input. These mechanical properties changes with temperature so their values are also considered as temperature dependent. Stress and strain fields are dependent on evolution of plastic strains so kinematic hardening in addition to Von Misses yield criteria is assumed which is valid for clad, interface and the substrate region"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000011_carpi.2012.6473371-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000011_carpi.2012.6473371-Figure12-1.png",
        "caption": "Fig. 12. Aerial inspection using UT and a manipulator system.",
        "texts": [
            " The need for laborious and dangerous installations of scaffolds, elevator systems or climbing utilities may be significantly reduced. Measurement strategies may vary from sampling wall thickness at individual points of a structure to full line or surface scans. Several approaches are feasible to implement aerial inspection by contact. The aerial vehicle may attach to the structure either by vectoring its thrust towards the inspected surface or using suitable adhesion technologies to dock e.g. such as [11]. The requirement of an inspection manipulation system as visualized in Figure 12 is also anticipated to ensure proper contact of the inspection probe and alleviate possible alignment issues when placing a probe on uneven surfaces such as boiler tubes. Based on the industrial and the generic inspection scenarios a set of requirements can be derived for future aerial inspection systems. This list of requirements is discussed in detail next and may serve as a benchmark to validate future aerial inspection robots within the presented application field. The first set of requirements targets the general system structure and constrains dimension, weight, payload and endurance of future aerial inspection robots: \u2022 The entire system must be robust and reliable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002315_ffe.12237-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002315_ffe.12237-Figure5-1.png",
        "caption": "Fig. 5 Detail of the half meshed model.",
        "texts": [
            " To simulate the balanced shaft, the cases corresponding to the same three crack lengths and eight rotation angles have been also modelled with e = 0 in order to compare with the corresponding unbalanced cases. For the FE analysis, the complete 3D model of the shaft has been considered because there is no symmetry in the problem. According to other authors,24 linear stress and strain distributions are assumed; the mesh of the three dimensional model is made by employing 8-nodes linear brick elements (C3D8R) with increasing density in the regions surrounding the crack (in Fig. 5a, for the case \u03b1 = 0.25, a detail of the meshed half model is shown). In order to determine the level of mesh refinement, convergence analysis was carried out for each crack length; the FE model finally included about 220 000 elements and 235 000 nodes. On the other hand, to calculate accurate SIF values, it is very important to avoid the interpenetration between the crack faces when the crack is in the compression zone (closed), so a surface-to-surface contact interaction has been defined. One of them is called \u2018master surface\u2019 and the other is called \u2018slave surface\u2019",
            " Hence, to complete the definition of the contact model, it is necessary to establish both normal and tangential properties between the crack faces. Regarding the normal contact properties, \u2018hard\u2019 contact option is used; this does not allow the penetration of the slave surface into the master surface at the constraint locations and prevents the transfer of tensile stress across the interface. The chosen tangential property corresponds to a \u2018rough\u2019 friction, it means an infinite coefficient of friction that avoids the relative sliding motion between two contacting surfaces (in Fig. 5b, one can see a detail of the surfaceto-surface contact interaction zone inside a translucent full model, \u03b1 = 0.25). The ABAQUS/Standard code allows several fracture mechanics parameters to be evaluated,39 such us the Jintegral or the SIF for modes I, II and III. In this work, an energetic approach has been chosen to determine the SIF distribution along the crack front. First of all, ABAQUS/Standard calculates the J-Integral values. This \u00a9 2014 Wiley Publishing Ltd. Fatigue Fract Engng Mater Struct, 2015, 38, 352\u2013367 integral characterizes the energy release associated with the crack growth as follows:45 J \u00bc \u222b \u0393 Udy t \u2202d \u2202x ds (3) where the \u0393 path encloses the crack tip (its initial and final points lie on the two crack faces)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002732_eitech.2016.7519641-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002732_eitech.2016.7519641-Figure5-1.png",
        "caption": "Fig. 5: Stator flux \u03c8s variation in space vector.",
        "texts": [
            "Ts (13) For a switch-on time Ts of the voltage Vs, (13) implies that the variation of the stator flux linkage will be in the direction of the applied voltage Vs in a period of time Ts. To control the magnitude of stator flux to be in the hysteresis band, it must select one of the non-zero voltage vectors. In each sector k, two adjacent voltage vectors Vk\u00b11 and Vk\u00b12 are selected to modify the amplitude of stator flux linkage. In this way, the stator flux can be controlled at the desired value by selecting the proper voltage as shown in fig.5. When the stator flux \u03c8s in sector k, the control of stator flux and torque can be achieved by selecting one of the eight vectors Vs, for a given k: \u2022 k+1: Te \u2197 , \u03c8s \u2197. \u2022 k+2: Te \u2197, \u03c8s \u2198. \u2022 k-1: Te \u2198 , \u03c8s \u2197. \u2022 k-2: Te \u2198 , \u03c8s \u2198. V0 and V7 are selected when the stator flux rotation should be stopped. Then, the torque will be decreased and the magnitude of stator flux remains unchanged. To ensure the quick dynamic torque response, it should control the rotating speed of the load angle \u03b4\u0307 as fast as possible, for this purpose the switching table is established"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003047_jmr.2017.174-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003047_jmr.2017.174-Figure2-1.png",
        "caption": "FIG. 2. Model of geometry and finite element.",
        "texts": [
            " University of Warwick, on 18 May 2017 at 06:20:36, subject to the Cambridge Core terms of use, available at where a and b are the dynamic coefficient of dendrites tip growth, m is the slope of the liquidus, C0 is the initial solute concentration, k is the partition coefficient, and D is the diffusion coefficient. By the calculation, for the industrial pure titanium TA1, a 5 1.64 10 6 and b 5 1.03 10 7. The temperature distribution has a significant effect on the grain microstructure during the continuous solidification. To obtain the temperature distribution before coupling the CA method, the solution must be carried out on the specific conditions including geometric conditions, initial conditions, and boundary conditions. (i) Figure 2 shows the model of geometry and finite element enmeshment. In consideration of symmetry reasons, only half of the geometry (625 210 8000 mm) is represented. To reduce the computation time and improve the calculation precision, the size of the surface mesh is 10 mm on the titanium slab ingot and the size of the surface mesh on the crystallizer is 15 mm. And then the volume mesh can be generated by dragging the surface mesh. The total number of elements is 1,315,867. Since the Lagrangian representation could have been used for both the FE and CA calculations, the cooling conditions associated with the crystallizer would have been translated along the lateral surface of the nonmoving slab ingot.17 With the boundary movement method calculating the temperature distribution, the initial position of the crystallizer, which moves up in the pulling speed assumed to be constant in this simulation, is in the bottom of the ingot shown in Fig. 2. The calculation zone, which is 800 mm distant from the bottom face of the ingot, generates automatically a regular network of square cells with 200 lm spacing. The calculation zone we choose is to simulate the microstructure evolution using CAFE method and compare with the experimental results of cross-section of the selection zone. (ii) The titanium molten pool surface in the crystallizer heated by the electron beam gun keeps the constant pouring temperature we assumed, which can be controlled by adjusting the power of the electron beam gun"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002067_s11249-013-0291-y-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002067_s11249-013-0291-y-Figure9-1.png",
        "caption": "Fig. 9 Structure Drawing of the friction torque sensor unit",
        "texts": [
            " The design idea and installation of the FTTD are explained as follows. (1) The EMLD (as indicated by item 10 in Fig. 7) is connected directly by screws with the stator seat. The stator is fixed in the stator seat, and they are in general considered as a unit. So, the friction force generated by the mating faces is transmitted directly to the EMLD. (2) A U-shaped fork is designed, and mounted on the EMLD (Fig. 8). (3) A friction torque sensor unit (as indicated by item 2 in Fig. 7) is designed, and its structure drawing is shown in Fig. 9. The photograph of this sensor unit is shown in Fig. 10. The strain gauge (as indicated by item 3 in Fig. 7) is fixed on the cantilever part of sensor unit\u2014see also Fig. 9. As shown in Fig. 9, the round end of the sensor unit is introduced in the U-shaped fork with no clearance. The friction torque sensor unit is mounted on the electromagnetic cover (as indicated by item 1 in Fig. 7) by two screws, and the electromagnetic cover is connected with the seal chamber. The friction force of the face seal is measured by the strain gauge. The calibration data between the strain signal and the force is shown in Table 1. The system calibration is completed in the standard force-strain test platform by the sensor company affiliated to the Aerospace Research Institute, and the whole calibration process simulates the actual stress conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001482_0954406218782285-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001482_0954406218782285-Figure6-1.png",
        "caption": "Figure 6. Comprehensive fault simulation test bench.",
        "texts": [
            " Table 1 gives the evaluation parameters of the de-noising effect of each method. Obviously, Wavelet soft threshold method and the presented method are both effective. By contrast, the proposed method has the better inhibitory effect to the noise. The presented method was used to analyze the actual bearing fault signal. A comprehensive fault simulation test bench is used to do gear fault diagnosis test. The test bench is composed of motor, mechanical transmission device, sensor, the hardware circuit, computer, and related software. Figure 6 shows the test bench. Inner diameter is 27.6679mm. Outer diameter is 57.1476mm. Thickness is 16.4518mm. Pitch diameter is 47.1532mm. Rolling body diameter is 9.4671mm. In outer ring, damage diameter is 0.1659mm which is made by the electric spark machining. Rolling element number is 7. The contact angle is 0. Bearing speed is 1830 r/min. The sampling frequency of vibration signal is 12 kHz. The number of sampling points is 3072. Figure 7 shows the time domain waveform of the measured inner race bearing fault signal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001956_citcon.2015.7122604-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001956_citcon.2015.7122604-Figure4-1.png",
        "caption": "Fig. 4: Shaft grounding currents. Note path to ground through the bearings in the motor and in the driven machine.",
        "texts": [
            "00 \u00a9 IEEE Page 2 of 11 Page 3 of 11 2014-CIC-1045 The third type is because the converter tries to emulate a sine wave supply by PWM signals, which have a high frequency switching and very fast switches (steep edge pulses). The contacts between the rolling elements and rings form a capacitor. The fast switching causes capacitive discharge currents. A similar problem to the poor cabling, the Protective Earth voltage is due to the common mode voltage from the inverter, and it is at a higher frequency. If the impedance of the return cable is too high, and if the stator grounding is poor, the current can take the path from the stator, through the bearings and through the shaft to ground in the driven machinery. Fig. 4. There is a common mode disturbance, causing current asymmetry between the three phases in the stator windings. The current sum over the stator circumference is not zero - HF flux variation is surrounding the shaft, creating a HF shaft voltage. Therefore, there is a risk for an axially flowing current through the rotor which runs through one bearing and back through the other. Fig. 5. In rolling element bearings, the rolling element is separated from the rings by a lubricant film. The lubricant film acts as a dielectric charged from the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000053_1464419314566086-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000053_1464419314566086-Figure1-1.png",
        "caption": "Figure 1. Structure of a spherical roller bearing. 1. Inner ring; 2. Cage; 3. Roller; 4. Outer ring.",
        "texts": [
            " The maximum loads of the rollers in the two rows neighboring on the off-sized roller/rollers can also be affected. Meanwhile, roller diameter error will largely affect the axis orbit and inner ring radial displacement of spherical roller bearings. Spherical roller bearing, off-sized rollers, load distribution, axis orbit Date received: 19 October 2014; accepted: 5 December 2014 Spherical roller bearings (SRBs), also known as selfaligning roller bearings, are widely used mechanical parts in low speed and heavy load industrial equipment, such as rolling mills1 and paper machines.2 As shown in Figure 1, a typical SRB structure usually consists of two rows of symmetric or asymmetric fully crowned rollers, a cage, an inner ring with two spherical raceways, and an outer ring with one spherical raceway. It is the special spherical structure of the rollers and raceways that provides the self-aligning function of the SRBs, i.e., a certain amount of inner ring tilting angle, misalignment, and shaft deflection are allowed in SRBs. It is difficult to analyze the mechanical performance of the SRB due to its relatively complicated structure",
            "comDownloaded from Ge \u00bc 7:3242 10 4k 0:2743e R 1=3ex \u00f013\u00de The total normal elastic deformation between the roller and the raceways can be expressed as n \u00bc i \u00fe e \u00bc Gi \u00fe Ge\u00f0 \u00deQ2=3 \u00bc GnQ 2=3 \u00f014\u00de where Gn \u00bc 7:3242 10 4\u00f0k 0:2743i R 1=3ix \u00fe k 0:2743e R 1=3ex \u00de \u00f015\u00de Then, the contact load due to elastic deformation can be expressed as29 Q \u00bc Kn 1:5 n \u00f016\u00de In equation (16), Kn is the stiffness coefficient, which can be calculated by Kn \u00bc G 1:5n \u00bc 5:045 104 k 0:2743i R 1=3ix \u00fe k 0:2743e R 1=3ex 1:5 \u00f017\u00de The normal loads of the jth roller Q1 j and Q2 j illustrated in Figure 1 can be expressed as Q1 j \u00bc Kn 1:5 1 j \u00f018\u00de Q2 j \u00bc Kn 1:5 2 j \u00f019\u00de The normal elastic deformation of roller raceway 1 j and 2 j are expressed as33 1 j \u00bc r cos 1j Pd 2 \u00f020\u00de 2 j \u00bc r cos 2j Pd 2 \u00f021\u00de where r is the inner ring radial shift and Pd is the bearing radial clearance. The half of load zone angle L can be calculated by the following equation L \u00bc arccos Pd 2 r \u00f022\u00de Figure 3 shows the roller load analysis in an aligned SRB. When the inner ring has a tilting angle, the loading direction of the bearing will also have a tilting angle "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000257_s00707-015-1333-3-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000257_s00707-015-1333-3-Figure1-1.png",
        "caption": "Fig. 1 Schematic model of the system",
        "texts": [
            " 3 gives an analytical solution of the pitch motion by using elliptical functions; Sect. 4 determines the Melnikov Z. Pang \u00b7 B. Yu \u00b7 D. Jin (B) State Key Laboratory of Mechanics and Control of Mechanical Structures, Nanjing University of Aeronautics and Astronautics, Nanjing 20016, China E-mail: jindp@nuaa.edu.cn Tel.: +86 25 84890251 Z. Pang E-mail: pangzj@nuaa.edu.cn B. Yu E-mail: yu_bensong@nuaa.edu.cn function of the perturbed spacecraft; and Sect. 5 gives an analytical criterion for preventing chaotic motion based on the Melnikov function. As shown in Fig. 1, consider a spacecraft which drags a satellite by a viscoelastic tether of unstrained length l. The spacecraft is treated as a rigid body of mass M1, while the satellite is envisioned to be a point of mass M2, with the assumption that M2 M1. The Earth-centered inertial frame is denoted by O-XYZ, the origin of which is located at the center of the Earth. The origin of orbital frame C-xyz is put at the center of mass of the system, with the x-axis pointing toward the center of the Earth, the y-axis following the tangent of orbit, and the z-axis completing the right-handed coordinate system",
            " The tension forces can be modeled by a linear Kelvin\u2013Voigt law of viscoelasticity in the following form: Pi\u22121,i = E A ( \u03b7i\u22121,i \u2212 1 + c\u03b7\u0307i\u22121,i ) (4) where E is Young\u2019s modulus of the tether, A the cross-sectional area of the tether, \u03b7i\u22121,i the elongation of the tether segment between node i \u2212 1 and node i, \u03b7\u0307i\u22121,i the elongation rate, and c the dissipation constant of the tether. In order to reveal the configurations of the elastic tether in the process of the motion of system, the in-plane pitch motion of the system is considered. The two pitch angles of the spacecraft and the tether are denoted by \u03b1 and \u03b8 , as indicated in Fig. 1. As a demonstration example, a set of parameters was taken as follows: the masses of spacecraft and satellite M1 = 1000 \u00d7 103 kg and M2 = 0.3 \u00d7 103 kg; the moments of inertia of the spacecraft Ixb = 20 \u00d7 106 kg m2, Iyb = 30 \u00d7 106 kg m2, Izb = 30 \u00d7 106 kg m2; the offset distance \u03c1 = 1 m; the normalized length of the tether l = 1 km; the linear density \u03bc = 1.7318 \u00d7 10\u22123 kg/m; Young\u2019s modulus E = 110 GPa; and the cross-sectional area A = 1.7671 \u00d7 10\u22126 m2. The dissipation constant of the tether was set as c = 0",
            " To get insight into the chaotic behaviors of the spacecraft, a simplified rod model for the tether will be employed. It is assumed that the mass of the spacecraft is much larger than that of the subsatellite, and the center of mass of the system coincides with that of the spacecraft, where the mother satellite moves in an unperturbed Kepler circle orbit of radius R and true anomaly \u03bd. The subsatellite is attached to the spacecraft by an inelastic massless tether of length l at a joint point of distance, \u03c1, to the mass center C of the spacecraft, \u03c1 l. According to Fig. 1, the position vectors of the spacecraft and the subsatellite in the inertial frame are, respectively, R1 = [ R cos v R sin v 0 ]T , (5) and R2 = \u23a1 \u23a3 cos(v)(R \u2212 l cos \u03b8 \u2212 \u03c1 cos \u03b1) \u2212 sin(v)(\u2212\u03c1 sin \u03b1 \u2212 l sin \u03b8) sin(v)(R \u2212 l cos \u03b8 \u2212 \u03c1 cos \u03b1) + cos(v)(\u2212\u03c1 sin \u03b1 \u2212 l sin \u03b8) 0 \u23a4 \u23a6 . (6) The potential energy of the system is in the following form [17]: V = \u2212\u03bce M2 |R2| \u2212 \u03bce M1 |R1| \u2212 1 2 \u03bce |R1|3 ( Ixb + Iyb + Izb ) + 3 2 \u03bce |R1|3 ( Ixb cos2 \u03b1 + Iyb sin2 \u03b1 + Izb ) . (7) The kinetic energy of the system is T = 1 2 Izb ( d\u03b1 dt + d\u03bd dt )2 + 1 2 M2 \u2223 \u2223 \u2223 \u2223 dR2 dt \u2223 \u2223 \u2223 \u2223 2 + 1 2 M1 \u2223 \u2223 \u2223 \u2223 dR1 dt \u2223 \u2223 \u2223 \u2223 2 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure13-1.png",
        "caption": "Fig. 13. Planar 3 R RR parallel robot, and rectangular geometry of each distal link A j B j .",
        "texts": [
            "1 , both the interior barriers and the external boundaries obtained when omitting collisions are altered when forbidding these collisions between different links. In the following subsection, we will show an example where forbidding collisions does not affect the external boundaries of the workspace (i.e., the shape of the workspace), but drastically modifies its interior barriers (i.e., its internal structure). In this subsection, the proposed method will be applied to obtain the barriers of the reachable workspace of the planar 3 R RR parallel robot shown in Fig. 13 . This robot can be considered as a redundant robot if we actuate angles ( \u03b11 , \u03b12 , \u03b13 ) to control the position ( x, y ) of the center of the end-effector (which is an equilateral triangle B 1 B 2 B 3 of side h ), without caring about its orientation \u03c6. In that case, self-motion manifolds are curves in the 3-dimensional joint space of actuated angles ( \u03b11 , \u03b12 , \u03b13 ). However, the method proposed in Section 3 cannot be directly applied to angular joint coordinates. This is because angles undergo wrapping [i",
            " , \u03b86 ] T defined as follows: \u03b81 = cos \u03b11 \u03b83 = cos \u03b12 \u03b85 = cos \u03b13 (13) \u03b82 = sin \u03b11 \u03b84 = sin \u03b12 \u03b86 = sin \u03b13 These \u201caugmented\u201d joint coordinates will be subject to the following additional constraints: \u03b82 1 + \u03b82 2 \u2212 1 = 0 (14) \u03b82 3 + \u03b82 4 \u2212 1 = 0 (15) \u03b82 5 + \u03b82 6 \u2212 1 = 0 (16) In this way, we have doubled the dimension of the original joint space (which is now 6-dimensional), but the wrapping problem is now avoided and the proposed method can be applied exactly as described in Section 3 . In this case, self-motion manifolds are curves in the 6-dimensional joint space (\u03b81 , . . . , \u03b86 ) . The task variables are the position coordinates t = [ x, y ] T of the end-effector, and the passive variable is angle \u03c8 = [ \u03c6] T . The following restrictions can be derived from Fig. 13 : \u2225\u2225\u2225\u2225 [ x y ] \u2212 a 1 [ \u03b81 \u03b82 ] \u2212 h \u221a 3 [ cos (\u03c6 + \u03c0 6 ) sin (\u03c6 + \u03c0 6 ) ]\u2225\u2225\u2225\u2225 2 \u2212 b 2 1 = 0 (17) \u2225\u2225\u2225\u2225 [ x y ] \u2212 [ c 2 x 0 ] \u2212 a 2 [ \u03b83 \u03b84 ] \u2212 h \u221a 3 [ cos (\u03c6 + 5 \u03c0 6 ) sin (\u03c6 + 5 \u03c0 6 ) ]\u2225\u2225\u2225\u2225 2 \u2212 b 2 2 = 0 (18) \u2225\u2225\u2225\u2225 [ x y ] \u2212 [ c 3 x c 3 y ] \u2212 a 3 [ \u03b85 \u03b86 ] \u2212 h \u221a 3 [ cos (\u03c6 \u2212 \u03c0 2 ) sin (\u03c6 \u2212 \u03c0 2 ) ]\u2225\u2225\u2225\u2225 2 \u2212 b 2 3 = 0 (19) where a j are the lengths of proximal links O j A j , b j are the lengths of distal links A j B j , and { c 2 x , c 3 x , c 3 y } determine the positions of joints O 2 and O 3 as illustrated in Fig. 13 . Eqs. (14) to (19) constitute the system (2) defining the self-motion manifolds in this particular example. Since in this example self-motion manifolds are 1-dimensional, Algorithm 2 reduces to sweeping each of the six axes \u03b8 j of the joint space independently, and calculating all five unknown joint coordinates (and passive angle \u03c6) in terms of the swept joint coordinate in each case. Note that joint coordinates in this example must be swept between \u03b8min j = \u22121 and \u03b8max j = 1 , since the \u201caugmented\u201d joint coordinates are sines or cosines",
            " Next, the IK problems of serial chains O 1 A 1 B 1 and O 3 A 3 B 3 are solved, obtaining ( \u03b11 , \u03b13 ). Finally, ( \u03b81 , \u03b82 , \u03b85 , \u03b86 ) are computed from Eq. (13) . It is easy to extend this procedure to the cases where \u03b85 or \u03b86 are swept. For the examples that will be illustrated next, we will consider that there are not joint limits and that distal links A 1 B 1 , A 2 B 2 , and A 3 B 3 move in the same plane and, therefore, their mechanical interference should be forbidden. For collision testing, we will consider that distal links have rectangular shape with the dimensions indicated in Fig. 13 : Their length is (b j + 2 \u03bb) , and their width is w . The collision test between a pair of distal links (rectangles) can be easily performed using the separating axis theorem. Next, we will apply the proposed method to an example of a 3 R RR robot with the following parameters: a j = 0 . 18 , b j = 0 . 22 , h = 0 . 15 , c 2 x = 0 . 5 , c 3 x = c 3 y = 0 . 3 , 2 \u03bb = w = 0 . 035 (this geometry precisely corresponds to the robot depicted in Fig. 13 ). The method described in Section 3 is applied to obtain the barriers of the reachable workspace of this robot in the following box of the task space: T = [0 , 0 . 5] \u00d7 [ \u22120 . 19 , 0 . 41] . The following parameters are used to execute the proposed method: N j s = N t i = 200 , \u03c3c = 1 . 5 , \u03c3m = 2 . The calculation of barriers in this example is again distributed over N p = 8 pro- cesses working in parallel. In this case, box T is divided into eight equal parts along the vertical y axis (as indicated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002491_s11433-014-5404-6-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002491_s11433-014-5404-6-Figure1-1.png",
        "caption": "Figure 1 Cross section of radial lip seal.",
        "texts": [
            " rotary lip seal, elastohydrodynamics, surface texture, pumping rate, friction torque PACS number(s): 47.11.Df, 47.15.Rq, 47.55.Ca, 47.85.Dh Citation: Guo F, Jia X H, Gao Z, et al. The effect of texture on the shaft surface on the sealing performance of radial lip seals. Sci China-Phys Mech Astron, 2014, 57: 13431351, doi: 10.1007/s11433-014-5404-6 Radial lip seals are widely used in the lubrication area of rotating machinery to prevent leakage and exclude contamination. A schematic of a typical installed lip seal including components is shown in Figure 1, and the sealing zone between the lip and the shaft is shown in Figure 2. In all successful lip seals, leakage is prevented by a reverse pumping action. This reverse pumping action leads to the lubricating oil flowing from the air-side of the lip seal to the oil-side, the direction of which reverses the normal leakage flow direction from the oil-side of lip seals to the air-side. The pumping action of lip seal is considered to be principally induced by the asperities on the lip surface. When the lip seal is in use, the shear stress between the lip and shaft makes the asperities produce vane-like patterns, then resulting in axial reverse pumping flow toward the oil side by the rotating shaft",
            " The first term hs is considered to be p s , in texture region, 0, in non-texture region,    h h (13) where hp represents the texture depth. The second term h is calculated by the deformation analysis, which utilizes the influence coefficient method. For discretized form, avg sc 1 \u2206 ( ) ( ) ,    m i ik k k h I p p (14) where pavg is the fluid pressure averaged over on cycle in the x direction. An off-line finite element structural analysis (FEA), ANSYS, is used to calculate the influence coefficient matrix I and the static contact pressure distribution psc for the lip seal model shown in Figure 1. The effect of texture on the shaft can be ignored in the finite element analysis, for the texture depth is in several micrometers range. Solid element \u201cplane183\u201d is axisymmetric by default, and \u201ctarge169\u201d and \u201cconta172\u201d are used to define contact pair. The mesh size is 4552 elements, which is selected according to a mesh refinement study. The seal is analyzed using a Mooney-Rivlin model (C10=1.1 MPa, C01=0.67 MPa, corresponding to the elastomer material), while the shaft is treated as elastic (Young\u2019s modulus equals 200 GPa, Poisson\u2019s ratio equals 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002660_978-3-319-06590-8_48-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002660_978-3-319-06590-8_48-Figure5-1.png",
        "caption": "Fig. 5 Nachi 6205-2NSE9 single-row deep-groove ball bearing showing 1.6 mm long seeded fault across outer race",
        "texts": [
            "94 mm diameter, with a bearing pitch diameter of Dp = 39 mm (see Fig. 2). To simulate a localised bearing fault, a notch of length lO = 1.6 mm was seeded on the bearing outer race using electric spark erosion. The notch was seeded across the entire race such that it was approximately 0.5 mm deep at the centre, sufficient to prevent the balls contacting the bottom of the notch. The fault was carefully positioned in the centre of the load zone for all the tests. The bearing with seeded fault is shown in Fig. 5. In order to develop an alternative spall size estimation methodology, all the fault signals over the speed range tested were observed closely, with a particular view to isolating the entry event. The first step in this process was to create an \u2018event window\u2019 for each observable fault pass occurrence in the signal, an example of which is shown in Fig. 6. The event window was determined by first applying conventional envelope analysis to the signal, a process which nowadays can largely be automated [1]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure13-1.png",
        "caption": "Fig. 13. Density changes for the case of the spherical body moving in the water\u2013 vapor flow.",
        "texts": [],
        "surrounding_texts": [
            "The drag coefficient is calculated by a classical formula [7]: Cd \u00bc 2 Fd q A U2 ; \u00f05\u00de where Fd \u2013 the drag force, obtained from simulation results, q \u2013 water density, U \u2013 mean velocity of a water flow, m/s, A \u2013 reference area, m2."
        ]
    },
    {
        "image_filename": "designv11_13_0001482_0954406218782285-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001482_0954406218782285-Figure1-1.png",
        "caption": "Figure 1. Simulated FM\u2013AM signal: (a) time series; (b) 3D time\u2013frequency spectrum of the simulated signal from the generalized S transform; (c) the corresponding 2D time\u2013frequency spectrum of the simulated signal.",
        "texts": [
            " y\u00bc \u00bd\u00f01\u00fe 0:3sin\u00f02 15t\u00de\u00de cos\u00bd\u00f02 30t\u00fe 0:5sin\u00f02 15t\u00de\u00de \u00fe2\u00bd\u00f01\u00fe 0:5sin\u00f02 10t\u00de\u00de cos\u00bd\u00f02 130t\u00fe 0:5sin\u00f02 10t\u00de\u00de \u00f010\u00de The signal is composed of two kinds of amplitude and frequency modulated signals. One of the center frequencies is 30Hz and the modulation frequency is 15Hz. Another center frequency is 130Hz and modulation frequency is 10Hz. The sampling frequency is 1000Hz and the time length is 1 s. The time domain waveform of the simulated signal and the time\u2013frequency spectrum from the generalized S transform are shown in Figure 1. Figure 1(b) and (c) illustrate the three-dimensional time\u2013 frequency spectrum and its corresponding twodimensional brightness spectrum, respectively. It clearly shows the energy change of two frequency component with frequency and time. The frequency conversion and the energy accumulation area are clear as daylight and spectrum has a very high time\u2013 frequency resolution. The simulated signal is added to the random noise (Signal to noise ratio (SNR)\u00bc 5 dB) as the background noise. After adding noise, the time domain waveform of the signal and its time\u2013frequency spectrum from the generalized S transform are shown in Figure 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002298_j.msea.2014.01.081-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002298_j.msea.2014.01.081-Figure1-1.png",
        "caption": "Fig. 1. A schematic of PTCAP (a) first half cycle, (b) second half cycle and (c) die parameters.",
        "texts": [
            " The SPD techniques suitable for producing ultrafine grained (UFG) and NS cylindrical tubes are tubular channel angular pressing (TCAP) [9,10], high-pressure tube twisting (HPTT) [11], accumulative spin bonding (ASB) [12] and parallel tubular channel angular pressing (PTCAP) [13]. The PTCAP method is a novel SPD process which has several advantages compared to the other methods [14]. One advantage is the better strain homogeneity through the tube thickness, and the second is it needs lower process loads [14]. The PTCAP process consists of two half cycles shown schematically in Fig. 1. In the first half cycle, the first punch presses the tube material into the gap between mandrel and die including two shear zones to increase the tube diameter to its maximum value. Then, the tube is pressed back using the second punch in the second half cycle, decreasing the tube diameter to its initial value. After N passes of the PTCAP Contents lists available at ScienceDirect journal homepage: www.elsevier.com/locate/msea Materials Science & Engineering A http://dx.doi.org/10.1016/j.msea.2014"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003973_icieam.2019.8743076-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003973_icieam.2019.8743076-Figure4-1.png",
        "caption": "Fig. 4. Temperature distribution inside the induction motor, at t = 8000 s, considering a faulty motor condition. The short-circuit resistance is set at 1.35 .",
        "texts": [
            " As depicted in Fig. 3, there is a consistent and uniform distribution of temperatures inside the motor, typical of a healthy operation scenario. As expected, the rotor is the warmer part of the motor, followed by the stator. The rotor temperature is indeed quite uniform. Inside the stator, the windings are slightly warmer than the rest of the stator elements. Meanwhile, it is also observed that the temperature of the stator iron smoothly decreases in the direction of the stator outer boundaries. Fig. 4 depicts the temperature distribution map inside the motor for the scenario in which a short-circuit fault between 6 turns of phase W is imposed. The short-circuit resistance (1.35 ) is adjusted so that the short-circuit current flowing through the short-circuit branch equals the motor rated current. When comparing the results of Fig. 4 with the healthy condition (Fig. 3), it is observed that there is a general increment of the motor temperature. All motor components experience an increment of temperature of up to 10 \u00b0C. The rotor remains as the warmer component of the motor. As a consequence of the short-circuit fault, the temperature distribution along the stator is not uniform any longer. The slots that host the faulty turns are 2 - 4 \u00b0C warmer than all the other stator slots. The neighboring slots also suffer a minor temperature increment",
            " If a thermal analysis to the external boundary of the motor stator was carried out, the effects of the fault would remain undetected from outsider Fig. 5 depicts the temperature distribution map inside the motor for the scenario in which a short-circuit fault between 6 turns of phase W is imposed, with a short-circuit resistance of 0.1 . 0.1 . As stated in Fig. 5, the increment of the short-circuit severity promotes a sharp increment of the motor temperature, extensive to all motor components. Unlike the previous fault scenario (refer to Fig. 4), the slots that host the short-circuited turns are warmer than the rotor. Moreover, the temperature distribution along the stator is far more unbalanced than in the previous fault scenario (refer to Fig. 4). The temperature gradient between the slots that contain the faulty turns and the surrounding stator elements is greatly amplified. As a result, the progressive decrement of temperature in the direction of the stator outer boundary disappears. In this case, a simple observation of the stator surface temperature would allow to clearly notice the thermal unbalance along the stator. Despite the severe increment of temperature occurring near the shortcircuited turns, the temperature of the warmest regions remains slightly below the critical temperature, at which the integrity of the windings\u2019 insulation is sustained (155 \u00b0C)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002749_j.ijnonlinmec.2016.08.007-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002749_j.ijnonlinmec.2016.08.007-Figure2-1.png",
        "caption": "Fig. 2. Geometry of sphere to orthogonal parallelepiped contact.",
        "texts": [
            " This means that \u2016 \u2212 \u2016 \u2264 +rc c r ,1 2 1 2 while the penetration depth is evaluated from the relation = ( + ) \u2212 \u2212d r cr c ,1 2 1 2 as shown in Fig. 1. Moreover, the contact point x is defined by the intersection of the line that connects the centers and an arbitrarily chosen sphere among the two. The common normal vector n defines the line of impact and its direction is defined from the contact point to the selected sphere's center. The intersection test is performed by computing the point in the parallelepiped (or box) which is closest to the sphere. Let the box be centered at the origin and aligned with the axes, as shown in Fig. 2. In addition, let the dimensions of the box in the coordinate system with Cartesian axes x, y, z be l l l2 , 2 , 2x y z, respectively. Also, the sphere's center is located at point c\u00bc(c1 , c2 , c3 ), while the point in the box closest to the sphere's center is given by = ( ( \u2212 ) ( \u2212 ) ( \u2212 ))x c l l c l l c l lclamp , , ,clamp , , ,clamp , , ,x x y y z z1 2 3 where \u23a7 \u23a8\u23aa \u23a9\u23aa ( ) = < >c a b a c a b c b c clamp , , , if , if , otherwise The sphere intersects the box if the point x is contained in the sphere ||x c||r r, where r is the sphere's radius"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001985_ecce.2015.7309702-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001985_ecce.2015.7309702-Figure4-1.png",
        "caption": "Fig. 4. End shield structure of original motor with open type regreasable rolling element bearing (motor B)",
        "texts": [
            " In this work, small-scale journal bearings were designed and built for low voltage motors for testing the detectability of journal bearing faults under controlled fault conditions. The objective was to test motors with journal bearings under varying degrees of clearance with low voltage motors that are representative of large motors. The feasibility of detecting increased clearance in journal bearings was investigated on a 380 V, 5.5 kW, 2 pole induction motor. The main idea is to replace the rolling element bearings of a commercial low voltage test motor with custom-designed and built journal bearings, as shown in Fig. 4-5. To simplify the design and fabrication process, a motor with open-type rolling element bearings, shown in Fig. 4, was used. Open type bearings can be re-greased by injecting grease through the inlet and channel, and the internal bearing cap prevents the grease from entering the motor [32]. The rolling element bearings were removed and replaced with precision-machined steel journal bearings with shaft-bearing clearance of 60, 75, 90 \u03bcm (motors B1-B3), as shown in Figs. 5 and 6(a). The grease inlet/outlet, and channel were used for supplying the lubricant (oil) into the journal bearing through gravity feed, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000250_icaci.2013.6748525-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000250_icaci.2013.6748525-Figure2-1.png",
        "caption": "Fig. 2. A glidering motion cycle of an underwater glider",
        "texts": [
            " (1 1 ) L ( sin 2 ex + + - S111 ex -- - -- - - --Tns3 V Tnsl Tns3 V Tnsl cos2 ex ) + --- , m's3 1 . - (l'vl - Tnmrg(TmrI cos e + Tmr3 S111 e) Is3 -Tnrbg(TrbI cose + Trb3 sine) + (Tns3 q, Um, where D=(KDO + KDex2)V2, L=(KLO + KLex)V2, M=(KMO + KMex + Kqq)V2. The definitions of variables in (1) are listed in Table I. Underwater gliders travel by gliding up-and-down and follow a sawtooth-like pattern in the longitudinal plane. The control objective is to design buoyancy pump rate (Um) and mass moving rate (Umrx) such that the glider follows a desired sawtooth pattern (see Fig. 2). III. MPC FORMULAT ION Rewrite (1) as an equivalent discrete-time model as follow: x(k + 1) = f(x(k)) + g(x(k))u(k), y(k) = Cx(k), (2) subject to constraints: Umin ::.; u(k) ::.; Umax, \ufffdUmin ::.; \ufffdu(k) ::.; \ufffdUmax' Xmin ::.; x(k) ::.; Xmax, Ymin ::.; y(k) ::.; Ymax, (3) where x = [Vex q e 'tiL Tmrx]T E \ufffd6 is the state vector, U = rUm Umrx]T E \ufffd2 is the input vector, Y = e E \ufffdI is the output vector, C E \ufffdI X \ufffd6 is the output matrix, f and g are nonlinear functions derived from (1), Umin ::.; Umax, \ufffdUmin ::"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001559_tsmc.2018.2866856-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001559_tsmc.2018.2866856-Figure11-1.png",
        "caption": "Fig. 11. Variable-length pendulum.",
        "texts": [
            " However, the chattering is the most serious by using control law 3 as highlighted in Fig. 10(b). The reason is that the switching gain is overestimated, i.e., G(t) is stabilized at an unnecessary large value, as demonstrated in Fig. 10(c). By comparing Figs. 7(b), 9(c), and 10(c), it is evident that control law 1 is able to achieve the smallest amplitude of the switching gain. A variable-length pendulum control is further investigated in this paper. This is considered a benchmark problem for testing the performances of HOSMC algorithms [51], [52] (Fig. 11). All motions in this scenario are limited to the vertical plane. A load of mass m moves without friction along the pendulum rod. The distance from O to m is equivalent to L(t), which is unmeasurable. An engine is attached with the rod transmitting a torque w. The objective is to track the reference signal xc by the angular coordinate x of the rod. The dynamic model of the pendulum system is x\u0308 = \u22122 L\u0307(t) L(t) x\u0307 \u2212 g L(t) sin x + 1 mL(t)2 w (56) where m = 1 kg and g = 9.8 m/s2 is the gravitational constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002870_j.procir.2016.10.051-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002870_j.procir.2016.10.051-Figure3-1.png",
        "caption": "Fig. 3. Experimental setup",
        "texts": [
            " Moreover, after the spectrum was corrected by interpolation correction method with a Hanning window as below: 1 165.6f Hz , 2 386.5f Hz , 1 1.0086A , 2 1.0079A It is clear that this result is close to the theoretical value. 4. Application 4.1. Experimental set-up To evaluate the effectiveness of the proposed approach, the bearing vibration data from the Case Western Reserve University are used in the experiments [21]. The overall instruments configuration of the experimental setup is shown in Fig. 3. Bearing system fault diagnosis experiment platform consists of a 2hp motor (left), a torque transducer (middle), a dynamometer (right) and other control electronics (not shown). The test bearings support the motor shaft and the data were collected from an accelerometer mounted on an induction motor housing at the drive-end bearing. In the experiment, the test bearing was used to study only included one kind of surface fault: the bearing was damaged on the inner and outer race. The type of bearing is deep groove ball bearing 6205- 2RS, 9 balls and the bearing pitch diameter 39"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002889_s40436-016-0158-1-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002889_s40436-016-0158-1-Figure1-1.png",
        "caption": "Fig. 1 The simplified model of the motorized spindle",
        "texts": [
            " The assumed load (K) of angular contact angle contact ball bearing is given in Ref. [22] when its elastic deformation is da, the assumed load can be calculated by K \u00bc Fd zCd ; \u00f024\u00de where Fd is the radial load generated by deformation (N), and Cd is the constant of deformation. Cd \u00bc 34 300 G0:36 D 1=2 b ; \u00f025\u00de where G is the raceway groove curvature center coefficient, G = fi ? fe-1; fi is the internal raceway groove curvature radius coefficient; fe is the external one. These parameters are related to the selection of bearing. Figure 1 shows a high speed motorized spindle with adjustable preload. The thermal-structure of FEM model should be properly simplified. The rotor and the stator can be simplified as thick cylinders with evenly distributed heat source. Squirrel-cage three-phase asynchronous motor is often used in motorized spindle motor. In the simplified model, the two ends of the stator are considered as copper and the rotor ends are aluminum, while the middle part of them is silicon steel. The rolling elements of the bearing can be equivalent to a ring with the same cross-sectional area"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003924_j.jsv.2019.05.043-Figure19-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003924_j.jsv.2019.05.043-Figure19-1.png",
        "caption": "Fig. 19. The milling tests setup.",
        "texts": [
            " As a result, the contact stiffness and damping remain unchanged for different workpieces, which means that all the parameter identification works just need to be done once with screws unchanged. That shows the proposed method is very robust. Although the above analysis is simple, it\u2019s very important to extend the identified parameters to different workpieces, which will save much time and greatly decrease workload. In order to verify the effectiveness of the proposed method on milling forces compensation, milling tests are carried out on a three-axis milling machine (VMC-V5) as seen in Fig. 19. The workpiece is 6061 aluminium alloy. Two types of milling tools are adopted with one tooth and three teeth. The parameters of milling tools are as following: high-speed steel material, diameter 10 mm, overhang 65 mm, helix angle 0\u25e6 (one tooth) and 45\u25e6 (three teeth). The milling parameters are as below: axial milling depth 0.4 mm, radial milling depth 3 mm, feed per tooth 0.05mm/tooth, spindle speed 3000rpm-8000 rpm per 1000 rpm. The measured and compensated milling forces based on the predicted FRF in the previous section are presented in Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001801_978-90-481-9707-1_108-Figure50.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001801_978-90-481-9707-1_108-Figure50.1-1.png",
        "caption": "Fig. 50.1 Schematic of quadrotor UAV dynamics",
        "texts": [
            " The quadrotor system with payload is treated as the rigid body motion of slung-load systems, and the effects caused by the pendulum-like behavior of the slung load are considered in the general simulation model. The modeling method discussed in this chapter is to develop a systematic analytical formulation for micro UAVs with slung payload, which can be utilized to improve the control performance of the UAVs. Firstly, the dynamic model for quadrotor is introduced briefly. Taking inspiration from the work addressing on the classical helicopter, a quadrotor UAV system is studied in this chapter (Castillo et al. 2005a; Min et al. 2011). As shown in Fig. 50.1, the fuselage dynamics and the coordinate system can be divided into an earth frame fEg and a body frame fBg, which are used to describe the relative motions between the two coordinate frames. In order to simplify the highly nonlinear factors in quadrotor UAV system (Min et al. 2011), the following assumptions can be considered in developing the mathematical model of the quadrotor UAV, which are made based on the slower speed and lower altitude: Assumption 1. The center of mass and the body frame origin are assumed to coincide"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001401_j.measurement.2018.01.031-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001401_j.measurement.2018.01.031-Figure7-1.png",
        "caption": "Fig. 7. The Schematic diagram of DCS-150 type power closed type gear test machine.",
        "texts": [
            " 3, PB can be represented as, = \u2217\u239b \u239d \u2217 + \u2217 \u239e \u23a0 P \u03bc W d n \u03c0 \u03bc W d n \u03c02 2 60 2 60B b b bore b b bore1 1 1 1 2 2 2 2 (14) In the paper, the radial load Wb applied to bearing is replaced by the radial force Fr applied to gear. Taking a helical gear as an example, the force analysis of helical gear tooth is shown in Fig. 6. The calculation of radial force Fr is represented in Eqs. (15) and (16) =F T d 2 t (15) =F F \u03b1 \u03b2tan /cosr t n (16) Substituting PT and PB with (10), we can obtain the gear meshing efficiency \u03b7M. The DCS-150 type power closed type gear test machine is adopted (Fig. 7). The DCS-150 type power closed type gear test machine uses electric loading. The reaction gearbox is lubricated by circulating oil and water-cooled with disk shaped copper coil. The test gearbox is lubricated by oil pool and the oil type is 120# extreme pressure gear oil. Natural heat dissipation is used in the experiment and test data are recorded after heat balance. The sign of heat balance is that the oil temperature is no longer climbing after the gear box runs 20min continuously. The environment temperature is 45 \u00b0C in the experiment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003924_j.jsv.2019.05.043-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003924_j.jsv.2019.05.043-Figure3-1.png",
        "caption": "Fig. 3. Schematic diagram of cutting forces measurement system.",
        "texts": [
            " Based on RCSA, receptance from F1 to 2b can be obtained as follows H2b1(\ud835\udf14) = h2b2b(\ud835\udf14)((h2a2a(\ud835\udf14) + h2b2b(\ud835\udf14) + 1\u2215(k + i\ud835\udf14c))\u22121h2a1(\ud835\udf14) (2) where the component and assembly coordinates are coincident; X1 = x1, X2a = x2a, X2b = x2b; F1 = f1 and (k + i\ud835\udf14c)(x2b \u2212 x2a) = \u2212 f2b; k and c are contact stiffness and damping, respectively; H2b1(\ud835\udf14) = X2b\u2215F1 is the direct assemble response between the force F1 and the response H2b1; h2b2b(\ud835\udf14) = x2b\u2215f2b is the direct component receptance between the force f2b and the response x2b; h2a2a(\ud835\udf14) = x2a\u2215f2a is the direct component receptance between the force f2a and the response x2a; h2a1(\ud835\udf14) = x2a\u2215f1 is the cross component receptance between the force f1 and the response x2a. The detailed introduction of RCSA can refer to the literature [18]. Here, we just introduce RCSA simply and don\u2019t explain this method in detail. Based on TPA and RCSA, this section will derive the assembly FRF between tool tip and the dynamometer output. Here, milling and Kistler 9129A table dynamometer are selected for derivation as seen in Fig. 3. According to the rule of RCSA, the cutting forces measurement system is divided into three substructures: workpiece, table dynamometer with screws and joint surface between two previous substructures as presented in Fig. 4. Then, some symbols are defined for subsequent derivation as follows: Fr ,s1(s1 = x, y, z) are the real cutting forces at the tool tip position in the x, y, z directions; xw , i ,s2(s2 = z, y, z) are the workpiece displacements at the ith screw positions in the x, y, z directions; xs , i ,s2(s2 = z, y, z) are the ith screw displacements in the x, y, z directions; Fs , i ,s2(s2 = x, y, z) are the transmission forces of the ith screw in the x, y, z directions; Fm ,s3(s3 = x, y, z) are the measured cutting forces by dynamometer in the x, y, z directions; ki and ci are the contact stiffness and damping at the ith screw position, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000518_1.4028062-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000518_1.4028062-Figure4-1.png",
        "caption": "Fig. 4 Schematic diagram of torque calculation on the disks",
        "texts": [
            " On the right side, the pressure of eleventh separator applied by the flange of the support knob is defined as N2. Fi means the normal force for squeeze-film flow between two disks. The friction force between outer gear of separator disk and inner spline in HVD is fi. The friction force between inner gear of friction disk and outer spline in center shaft is fi\u00fe1. n is the number for the interfaces of multidisk friction pairs. The schematic of frictional torque Ti transmitted to any disk is shown in Fig. 4 and Ti can be expressed as follows: Ti \u00bc \u00f0b a lfricFi A rdA \u00bc \u00f0b a lfricFi A 2pr2dr \u00bc 2plfricFi 3A \u00f0b3 a3\u00de (10) After rearranging the solution of Eq. (10), the normal force Fi can be expressed as follows: Fi \u00bc 3TiA 2plfric b3 a3\u00f0 \u00de (11) Additionally, the friction torque Ti and the friction force fi, which included both inner and outer splines, must be balanced. Hence, the friction force can be derived by torque equilibrium as follows: fi \u00bc l0 2Ti di cos a (12) where a is the profile angle of spline pairs, l0 is the friction coefficient of spline pairs, l0 \u00bc 0:12, di is the pitch diameter of spline pairs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000195_tmag.2013.2245878-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000195_tmag.2013.2245878-Figure3-1.png",
        "caption": "Fig. 3. Self-alignment bearing using in PM-BLDC motor operation [7].",
        "texts": [
            " In previous work, most of the research was focused on dynamic UMP analysis [3]\u2013[5]; some work covers the static UMP [6], but no detailed research exists for the IUMP. In this paper, the detailed analysis will be concentrated on the incline UMP (IUMP). Normally, to improve the motor efficiency and power density, the motor air-gap should be as small as possible. However, the smaller the air-gap is, the more serious the rotor eccentricity will become, and thus the stronger the UMP will be. Besides HDD application, many PMSMs use self-alignment bearing [7], shown in Fig. 3, for reducing the rotor-bending force. It is clear that this kind of bearing could worsen the misalignment of the rotor, and consequently stronger IUMP could be induced. There are many kinds of PMSMmotors [8], [9]. In this paper, the analysis will be focused on the motors with surface PM ring Manuscript received November 30, 2012; revised February 01, 2013; accepted February 01, 2013. Date of current versionMay 30, 2013. Corresponding author: Y. Yu (e-mail: Yu_YinQuan@dsi.a-star.edu.sg). Color versions of one or more of the figures in this paper are available online at http://ieeexplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000481_jfm.2014.666-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000481_jfm.2014.666-Figure2-1.png",
        "caption": "FIGURE 2. Streamlines of the perturbation to the core flow, outside the concentration boundary layer, given by (2.27) and (2.30), for \u03b80 = 0 (no sedimentation of individual cells). (a) Re= 0, (b) Re= 20, (c) Re= 200.",
        "texts": [
            "32) The angle \u03b80 is given by (2.17), and represents the direction of the average velocity of the cells in the core of the cylinder, due to both swimming and sedimentation, relative to the basic solid-body rotation. However, its value does not affect the structure of the cell concentration distribution (2.16) or the resulting perturbation to the flow, given by (2.20) and (2.30). Hence, we plot the perturbation flow for the case \u03b80 = 0; the flow pattern is merely rotated for other values of \u03b80. Streamlines are plotted in figure 2 for Re = 0, 20 and 200, and in figure 3 we plot profiles of the tangential velocity outside the concentration boundary layer, given by (2.26a,b), (2.27) and (2.30), for the same values of Re. The predicted boundary-layer structure can be seen at the largest value of Re. It should be remembered that the plotted flow is a perturbation from the background solid-body rotation, so could only be directly observed if one used a co-rotating camera. From a laboratory-based perspective only the radial component of this motion is likely to be directly visible, although tangential velocity measurements could be made to check the predictions quantitatively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002286_j.triboint.2014.02.014-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002286_j.triboint.2014.02.014-Figure3-1.png",
        "caption": "Fig. 3. Piston secondary motion setup: (a) array of laser displacement sensors; (b) machining profile on piston crown.",
        "texts": [
            " The overlapping region (shown as the area with cross-hatch) indicates that the majority of the time the lubricating condition of the experimental rig running at 100 rpm to 500 rpm at 30 1C to 40 1C matches the lubricating condition of the actual engine running at the actual temperature and speed range of 1000 rpm to 5000 rpm. As reported by previous researchers, the piston motion can be classified into four components: translational motion along the cylinder bore axis, lateral motion, piston rotational motion around the piston pin axis, and a combination of piston lateral motion and rotational motion when the piston reciprocates [36\u201339]. These motions were captured using three laser displacement sensors, Fig. 3a. A flat slot and 451 profile were machined on the surface of the piston crown, as shown in Fig. 3b. Two laser beams were directed to the flat profile to capture the rotational motion of the piston obtained from the relative motion between the two laser beams, while the 451 profile was used to calculate the piston lateral motion with the assistance of the third laser beam. The distinct modes of the piston motion were classified based on the trigonometry algorithm. Details on the working principle and a description of the motion classification are presented in [40]. The contact damping and stiffness of the piston assembly are obtained from mobility measurement and this technique has been used to measure the piston dynamic properties in the study of piston slap [43]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000049_robio.2013.6739433-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000049_robio.2013.6739433-Figure1-1.png",
        "caption": "Fig. 1. Mechanical configuration of the turtle-like vehicle.",
        "texts": [
            " In Section III, we describe the CPG-based control in detail. The experimental results are shown in Section IV. Finally, Section V concludes the paper and summarizes the future work. A. Mechanical structure Focus on the design criterion of practicability, a turtle-like underwater vehicle equipped with four mechanical flippers has been developed by imitating the morphology of a marine turtle. The vehicle consists of a conical head, a cylindrical main body, two balancing cylinders, and four flipper actuator modules with wing-shaped foils attached. Fig. 1 shows a photograph of the turtle-like vehicle\u2019s mechanical configuration. The conical head of the vehicle acts as the control cabin, which contains the navigation module and the control boards. The cylindrical main body is the energy depot, which contains 24V 20AhLithium iron phosphate battery and the servo drivers of the flippers. The two balancing cylinders are used to adjust the position of the center of the gravity, and each cylinder consists of a thin cylindrical housing, a guide screw which attached with a clump weight, a servo motor and its driver"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000954_jsen.2017.2649686-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000954_jsen.2017.2649686-Figure1-1.png",
        "caption": "Fig. 1: a) Arrangement of the dialyzer bag containing the immobilized enzyme b) snapshot of the combined electrode as a CO2 biosensor.",
        "texts": [
            "5 grams of calcium chloride was dissolved into 500 ml of distilled water and kept at 4 0 C. 50 mg of sodium alginate was added into 5 ml of enzyme solution containing 5 mg of lyophilized enzyme in 0.02 M sodium phosphate buffer. The solution was added into 100 ml of cold calcium chloride solution. The calcium alginate beads, so formed, were washed with the buffer solution. The immobilized carbonic anhydrase enzyme was put in a dialyzing sac and the assembly was attached firmly at the tip of the sensing bulb of a combined pH-electrode using an Oring (Fig.1a). The enzymatic biosensor (Fig.1b) was then ready to measure the electrical potential, developed across the combination electrode in response to the concentration of the dissolved CO2 at 0 0 C. After the development of a biosensor based system, it is very important to assess its performance through examining various performance indices like sensitivity, stability, selectivity, shelf life, response time etc. The present work investigated a few of the performance indices to justify its clinical application. CO2 gas was bubbled for sixty minutes into a 500ml of double-distilled water at 0 0 C kept in a thoroughly washed 1000 ml conical flask, placed in an ice-water bath, at normal 1530-437X (c) 2016 IEEE"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001545_jfpe.12812-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001545_jfpe.12812-Figure2-1.png",
        "caption": "FIGURE 2 Surface response of acetic acid conversion for the synthesis of isoamyl acetate by mechanical agitation (a) and ultrasonic agitation (b). Ac: acetic acid; Al: isoamyl alcohol",
        "texts": [
            " According to Azudin et al. (2013), this hindering favors the product solubility and mass transfer in the reaction system. Equation 2 presents the second-order coded model, which describes the acetic acid conversion as a function of temperature (T, 8C) and acetic acid : isoamyl alcohol molar ratio (R, dimensionless). Within the range of values for the parameters evaluated, the correlation coefficient was 0.91 with Fcal (10.36) > Ftab (5.05), which allowed the construction of surface response presented in Figure 2a. Conversion \u00f0%\u00de584:74116:493R29:193R2116:863T 211:243T214:253T3R (2) The highest acetic acid conversions for the synthesis of isoamyl acetate were achieved in the temperature range of 56\u201361 8C and acetic acid : isoamyl alcohol molar ratio of 1:8\u20131:9, exceeding 90% of conversion (molar basis). These values are comparable to findings reported elsewhere. For example, the conversion of butyric acid for the synthesis of isoamyl butyrate was 83% during 600 min using immobilized Novozym 435 (Bansode & Rathod, 2014) and 73% during 480 min using T",
            " (2015) indicate the higher conversions as a consequence of cavitation bubbles that increase the solubility of molecules. Therefore, it increases the reaction rates using low energy. Equation 3 presents the second-order coded model, which describes the acetic acid conversion as a function of temperature (T, 8C), acetic acid : isoamyl alcohol molar ratio (R, dimensionless), and ultrasonic power (P, W). Within the range of values for the parameters evaluated, the correlation coefficient was 0.97 with Fcal (34.95) > Ftab (3.68), which allowed the construction of surface response presented in Figure 2b. Conversion \u00f0%\u00de593:61112:393T29:733T217:473R 24:473R217:543P24:183P223:403T3R22:783T3P (3) The highest acetic acid conversions for the synthesis of isoamyl acetate by ultrasonic agitation were achieved in the temperature range of 52\u201370 8C, acetic acid : isoamyl alcohol molar ratio of 1:6.8\u20131:9.8 and ultrasonic power 79\u2013110 W, exceeding 90% of conversion (molar basis). There is a trend of increasing the conversions by using values of the parameters near to the central point (50 8C, Ac : Al 1:7, and 79 W) especially in the region of temperature, molar ratio of substrates and ultrasonic power a few superior of the central point. The interaction between the parameters (Figure 2b) provides an indication that their highest values tested in the assays are not the best ones because the conversions start decreasing in such extremities. Ultrasonic systems have been studied for esterification reactions as well. For example, approximately 94% of conversions of substrates were obtained for the synthesis of butyl butyrate during 720 min using lipase from T. lanuginosus (Lipozyme TLIM) immobilized on styrene\u2013 divinylbenzene beads (Martins et al., 2013). Likewise, 96% of conversion of substrates for biodiesel production during 180 min using Lipozyme TLIM was obtained (Subhedar & Gogate, 2016)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002808_s11071-016-3072-y-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002808_s11071-016-3072-y-Figure3-1.png",
        "caption": "Fig. 3 Simplified model of the manipulator",
        "texts": [
            " 1, oscillations of the tank vehicle on various landforms mainly involve three kinds of movement: (1) shake movement (linear movement along the axis yb); (2) pitch movement (rotational movement about the axis xb); and (3) roll movement (rotational movement about the axis zb),where the Obxbyb is the space coordinate attached to the tank vehicle (which is not shown in the figure). For simplicity, this paper focused on pitch movement only as an initial study. Ignoring the dynamics of the chain driven system, the above manipulator is simplified as shown in Fig. 3, where XOY denotes the inertia frame, whose Y axis is parallel with the direction of gravity; B1, B2 and B3 denote the mounted base, lifting device and transfer device (including payload), respectively; xoy denotes the frame attached to the base,whose origino is fixedon O;C1,C2 andC3 denote the centroid of B1, B2 and B3, respectively; yr2 is the position of the lifting device, \u03b81 and \u03b82 are the position angles of the mounted base and transfer device, respectively; L1 and L2 are geometric parameters, as defined in Fig. 3. The dynamic equation of motion can be derived using Lagrange equations of the second kind, assuming the mounted base as an extra link. Regarding the bases term of the dynamic equation as disturbance forces, then extract the manipulators term, we could obtain the uncertain dynamic model of the system as: H(q)q\u0308 + C(q, q\u0307)q\u0307 + G(q) = U + S (1) where q = [yr2,\u03b83] is the vector of generalized coordinates; H(q) is the inertia matrix of the system, which is asymmetric positive definite matrix; C(q, q\u0307)q\u0307 is the vector of the Coriolis/centripetal forces; G(q) denotes the gravity forces; S denotes the uncertain external force due to the oscillation of the base; and U is the control force exerting on the manipulator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003206_s0018151x1705008x-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003206_s0018151x1705008x-Figure1-1.png",
        "caption": "Fig. 1. To the statement of the problem of layer-by-layer sintering of a thin vertical wall (the laser-beam path of motion over the sintered-object surface is shown schematically): (a) general view of the operating region (prepared wall on the build platform and surrounding unsintered powder shown semitransparent) and (b) finite-element model for solving the problem within the plane statement (cross section of the sintered wall is shown).",
        "texts": [
            " 5 2017 ON THE POSSIBILITY OF STEADY-STATE SOLUTIONS APPLICATION 733 Let us consider a planar statement of the problem on the assumption that the sintered vertical wall is fairly extended (in the direction \u201cout of plane\u201d) and heat transfer at its distant faces can be neglected and accepting that an identical thermal mode is implemented in each cross section oriented perpendicular to the heat source motion direction. In other words, identical thermal cycles occur in each cross section of this kind. This process of layer-by-layer synthesis of a thin vertical wall is shown in Fig. 1. As a result of the calculation, one should determine the thermal state of the operating region, which includes the object, surrounding unsintered powder, and build platform. To this end, we solve the transient heat conduction problem at each step corresponding to a new deposited layer [11, 12]. Then, for a layer-by-layer growing body, we arrive at periodic local heating of its surface due to the laser action. Shrinkage of the powder is neglected in the first approximation. The shrinkage and corresponding change in the equivalent properties of a powder layer can be taken into account based on the thermokinetic approach proposed in [13]",
            " This domain consists of the operating platform \u03a9P = {0 \u2264 x \u2264 L, \u2013H \u2264 y \u2264 0} and the growing region \u03a9L = {0 \u2264 x \u2264 L, 0 \u2264 y \u2264 ih}, which includes the sintered body and surrounding unsintered powder. As a result, piecewise constant properties corresponding to the materials of the platform, sintered body, and powder are set in the \u03a9i domain. In the example under consideration, the wall is at the center of the calculation domain and has thickness d. The calculation will be performed for a half of the domain, taking into account that the problem is symmetric (Fig. 1b); the coordinate system used in the calculations is also shown in this figure. The height of the domain \u03a9i increases from above by the thickness of one layer h at each ith calculation step. In the region obtained, the transient heat conduction problem is solved in the time interval : (1) where Ti (x, y, t) is the temperature field, which is implemented in the calculation domain \u03a9i at the ith calculation step; is the heat f lux vector; and k(x, y, Ti), Cp(x, y, Ti), and p(x, y) are, respectively, thermal conductivity, specific heat, and density, the values of which are specified piecewise constant functions of coordinates and continuous functions of temperature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002286_j.triboint.2014.02.014-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002286_j.triboint.2014.02.014-Figure1-1.png",
        "caption": "Fig. 1. Experiment setup for piston friction force measurement.",
        "texts": [
            " In this study, the instantaneous friction force contributed by the piston skirt of a single-cylinder motorcycle engine was derived from the measured piston secondary motion and total friction force of the piston. The total piston friction comprises the piston ring friction and piston skirt friction. The instantaneous piston skirt friction can be determined from the difference between the measured total piston friction and piston ring friction force; the intermittent contact and transient nature of the piston skirt must be considered. Fig. 1 shows the assembly of the experimental setup, where an AC motor drives the crankshaft of the piston assembly via a pulley system with the motor speed controlled by a variable frequency controller. The cylinder block of a single-cylinder engine was machined and mounted on top of a dynamometer (Kistler Type 9272). The dynamometer measured the total friction force of the piston acting on the cylinder block. The signal from the dynamometer was amplified by a multichannel charge amplifier (Kistler Type 5070) and captured by the data acquisition system (imc CS3008-1)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000334_tac.2013.2274707-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000334_tac.2013.2274707-Figure3-1.png",
        "caption": "Fig. 3. Reach controllability in two 2-D examples.",
        "texts": [
            " By Proposition 1 this is a contradiction. Example 2: We illustrate the concept of reach controllability with a 2-D example. However, it must be noted that a true example can only be exhibited in dimension 4 and higher, since in dimensions 2 and 3, no system is not reach controllable while also satisfying the two necessary conditions, Proposition 3.1 of [7] and Theorem 1. This aspect will be further explored elsewhere. Here we illustrate a case when a 2-D example simultaneously fails reach controllability and Theorem 1. Consider Fig. 3(a). The velocity vectors at , produce an equilibrium . Adding a positive component to or results in a violation of the invariance conditions. The only option is to add to or . This in turn results in velocity vectors at and as depicted by dashed arrows. Clearly, the zero vector is in the convex hull of these two vectors so there will be an equilibrium in along segment . Therefore, RCP is not solvable. Notice in this example an equilibrium can appear in the interior of , apparently violating Theorem 2. This is because , so Theorem 2 actually does not apply. On the other hand, Fig. 3(b) shows an example where the system is reach controllable. Here and so it can be added to both and . This results in new velocity vectors at and depicted as dashed arrows. Clearly, the equilibrium is pushed out of the convex hull of these two points\u2014it now lies below the simplex. In the next result we relate reach controllability to the existence of a coordinate transformation that decomposes the dynamics into those that contribute to open-loop equilibria and quotient dynamics. It is noted that a geometric characterization of reach controllability has not yet been obtained, but the following result gives a first evidence that one may exist"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000739_s0263574713000829-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000739_s0263574713000829-Figure3-1.png",
        "caption": "Fig. 3. (Colour online) Screen shot of the robot recovering from a push with 800 N for 0.1 s.",
        "texts": [
            " This method is suitable to impulsive push, so all pushes in the experiments last for 0.1 s. The robot stands in place and keeps its torso upright. It has the same mass distribution as the real robot. The total mass of the robot is 96.9 kg. The torso has more than 70% of the total mass. So the http://journals.cambridge.org Downloaded: 28 Jun 2014 IP address: 155.198.30.43 torso represents the CoM approximately. The length of the forefoot is 0.182 m. The height of the CoM is 0.883 m. It is lower than in ref. [5] because our model squats down about 0.07 m initially. Figure 3 shows the screen shots of the robot recovering from a disturbance. At moment 0.1 s, apply a force with 800 N at the pelvis. Then, the robot deviates from its original position and velocity. Our method enables the robot to move its whole body to recover to its original position and attitude. In order to study the influence of different forces on torso motion, we do a group of simulations with 400, 600 and 800 N applied at the pelvis. The motion in the lateral plane influences the motion in the sagittal plane, but not much, so Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000166_s12239-012-0117-1-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000166_s12239-012-0117-1-Figure1-1.png",
        "caption": "Figure 1. Diagram of an electric power steering system.",
        "texts": [
            " KEY WORDS : EPS, SPMSM, Current sensorless, Fault tolerance In a vehicle, electric motor drive systems have begun replacing traditional hydraulic systems, due in part to the increasing demand for safer, more efficient and more comfortable vehicles. Specifically, EPS systems, which reduce the driver\u2019s steering effort by using the torque generated by an electric motor, have become quite popular. An SPMSM type of motor is most common as an actuator due to factors such as high power density, efficiency, and wide operating speed range (Liao and Du, 2003). Figure 1 shows the components of EPS and its interface. When a driver turns the steering wheel, the steering torque sensor generates the command signals with respect to the torsional angle between the input and output shaft. The steering commands are transmitted to the current controller, which controls and drives the motor for steering assistance (Choi et al., 2007; Lee et al., 2011b). As shown in Figure 1, the electric motor drive system for EPS includes various sensors, such as a current sensor, a motor position sensor and a steering angle sensor for higher precision control. One of the areas of research for safety frameworks of EPS safety is the mitigation strategies of LOA (Loss of Assist). LOA means that EPS cannot support the desired level of torque assistance to the driver due to sudden faults in the system. Most vehicle manufactures specify the safety criteria of the reduced assist torque when an EPS has some faults"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002146_j.mechmachtheory.2014.04.012-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002146_j.mechmachtheory.2014.04.012-Figure3-1.png",
        "caption": "Fig. 3. Shaving machine with an auxiliary crowning mechanism.",
        "texts": [
            " (1)\u2013(5), the locus and unit normal of the right-hand side surface of themodified rack cutter can bewritten as rs \u00bc xs u1; v1;\u03c61\u00f0 \u00de; ys u1; v1;\u03c61\u00f0 \u00de; zs u1; v1;\u03c61\u00f0 \u00de;1\u00bd T ; \u00f06\u00de ns \u00bc nxs u1; v1;\u03c61\u00f0 \u00de;nys u1; v1;\u03c61\u00f0 \u00de;nzs u1; v1;\u03c61\u00f0 \u00de h iT ; \u00f07\u00de where and v1 \u00bc csc\u03b2o1 \u22122du1ro1\u03c61 cos\u03b1on \u00fe u1 cos\u03b2o1 \u00fe 2d2u3 1 cos\u03b2o1 \u00fe ro1\u03c61 sin\u03b1on 2du1 cos\u03b1on\u2212 sin\u03b1on ; \u00f08\u00de xs \u00bc ro1 \u00fe u1 cos\u03b1on \u00fe du2 1 sin\u03b1on cos\u03c61 \u00fe ro1\u03c61 \u00fe u1 cos\u03b2o1 sin\u03b1on\u2212du2 1 cos\u03b2o1 cos\u03b1on\u2212v1 sin\u03b2o1 sin\u03c61; \u00f09\u00de ys \u00bc ro1 \u00fe u1 cos\u03b1on \u00fe du2 1 sin\u03b1on sin\u03c61 \u00fe \u2212ro1\u03c61\u2212u1 cos\u03b2o1 sin\u03b1on \u00fe du2 1 cos\u03b2o1 cos\u03b1on \u00fe v1 sin\u03b2o1 cos\u03c61; \u00f010\u00de zs \u00bc v1 cos\u03b2o1 \u00fe u1 sin\u03b2o1 sin\u03b1on\u2212du2 1 sin\u03b2o1 cos\u03b1on; \u00f011\u00de nxs \u00bc \u22122du1 cos\u03b1on \u00fe sin\u03b1on\u00f0 \u00de cos\u03c61\u2212 cos\u03b2o1 cos\u03b1on \u00fe 2du1 sin\u03b1on\u00f0 \u00de sin\u03c61; \u00f012\u00de nys \u00bc cos\u03b2o1 cos\u03b1on \u00fe 2du1 sin\u03b1on\u00f0 \u00de cos\u03c61 \u00fe \u22122du1 cos\u03b1on \u00fe sin\u03b1on\u00f0 \u00de sin\u03c61; \u00f013\u00de nzs \u00bc \u2212 cos\u03b1on \u00fe 2du1 sin\u03b1on\u00f0 \u00de sin\u03b2o1: \u00f014\u00de Considering Eqs. (6), (7), and (8) simultaneously yields the right-hand side tooth profile and the unit normal vectors of the modified involute helical shaving cutter. Longitudinal crowning of the work gear tooth flank is achieved by using a proposed crowning mechanism with an auxiliary rocking motion, as shown in Fig. 3, and the parameters of this auxiliary crowning mechanism are as shown in Fig. 4. The traveling distance dp of the pin along the guideway is converted from the traverse distance zt of the pivot and it can be expressed as follows: dp \u00bc dh \u00fe zt\u00f0 \u00de cos\u03b8\u00fe dv sin\u03b8\u2212 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi dh \u00fe zt\u00f0 \u00de cos\u03b8\u00fe dv sin\u03b8\u00bd 2\u2212zt 2dh \u00fe zt\u00f0 \u00de q ; \u00f015\u00de parameters dv and dh are the vertical and horizontal distances between the pin and the pivot, respectively, and \u03b8 is the where guideway inclination angle with respect to the horizontal slide (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002997_techno-ocean.2016.7890667-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002997_techno-ocean.2016.7890667-Figure2-1.png",
        "caption": "Fig. 2. Design of the semi-submersible ASV",
        "texts": [
            " The positioning of AUVs from ASV is established by Ultra Short Base Line (USBL) system, and the communication between AUVs and ASV is established by underwater acoustic communication. Therefore, the ASV is a relay system of air communication and underwater communication, and it should have an aerial part and an underwater part. 978-1-5090-2445-2/16/$31.00 \u00a92016 IEEE Techno-Ocean 2016 309 III. SEMI-SUBMERSIBLE ASV This semi-submersible ASV is similar to the cruising type AUV in shape except antenna part as shown in Fig.2. The ASV has an antenna named \"sword\" through the sea surface. The ASV position data is obtained by GPS antenna on top of the sword. The communication between the ASV and the research vessel is also established by wireless LAN antenna or Iridium satellite communication antenna on top of the sword. The draft of the semi-submersible ASV will be adjusted with approximately 1.5 m, and the sea surface level is kept in a range of the sword length. The components of the ASV is similar to the typical cruising type AUV, except the relay part of the positioning and communication in between the research vessel and multiple AUVs as shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002889_s40436-016-0158-1-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002889_s40436-016-0158-1-Figure4-1.png",
        "caption": "Fig. 4 The thermal deformation of the motorized spindle",
        "texts": [
            " To obtain the thermal characteristic of the simplified model, the ANSYS Workbench 15.0 commercial code is applied. The environment temperature of the simulation is 24 C and the rotating speed is 20 000 r/min. The internal temperature of the motorized spindle is shown in Fig. 3. As can be seen from Fig. 3, the main heat sources of the motorized spindle are the rotor, the stator, and the bearings. The front cooling water passage takes away a portion of the heat thus the surrounding temperature is relatively low. The thermal deformation of the motorized spindle is shown in Fig. 4. It shows that the thermal deformation of the motorized spindle begins from the center of the spindle. The deformation is gradually accumulated, and the deformation of the stator is larger. Front and rear bearings are the parts of high temperature and large deformation in the motorized spindle. The thermal deformation of the ends of the motorized spindle is larger due to the deformation accumulation. The cooling water passage of the motor takes away a portion of the heat, thus the thermal deformation is relatively small"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001934_1.4030344-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001934_1.4030344-Figure5-1.png",
        "caption": "Fig. 5 Sector magnets: (a) front view and (b) sectional side view",
        "texts": [
            " Since it is difficult to represent a cuboidal part in cylindrical coordinates, an attempt is made to replace a cuboidal magnet by sector magnets of equal surface area. To validate the assumption, a comparison has been made between the vertical load carrying capacities of square and sector magnets. The vertical force using Coulombian model for the sector magnets and square magnets is given below. Equation (1) can be used for estimating the force between two sector magnets in vertical (Y) direction (Fy,s) (as shown in Fig. 5) by replacing \u201ch1\u201d by a/2, \u201ch2\u201d by a/2, \u201ch3\u201d by b/2, and \u201ch4\u201d by b/2 in Eq. (2). The modified vertical force equation for sector magnets (Fy,s) is given as Fy;s \u00bc r1r2 4pl0 R za\u00f0 \u00de \u00fe R za \u00fe H B\u00f0 \u00de \u00fe R za \u00fe H\u00f0 \u00de \u00fe R za B\u00f0 \u00de\u00f0 \u00de (3) where R a\u00f0 \u00de \u00bc \u00f0a=2 a=2 \u00f0b=2 b=2 \u00f0R4 R3 \u00f0R2 R1 e\u00fe r12 cos h\u00f0 \u00de r34 cos h0\u00f0 \u00de\u00f0 \u00der12r34 r2 12 \u00fe r2 34 \u00fe e2 2r12r34 cos\u00f0h h0\u00de \u00fe 2e r12 cos h\u00f0 \u00de r34 cos h0\u00f0 \u00de\u00f0 \u00de \u00fe \u00f0a\u00de2 1:5 dr12dr34dhdh0 Two cubical magnets of sides a, b, and c and a0, b0, and c0 are shown in Fig. 6. za is the axial offset between the magnets"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003358_s11837-018-2794-3-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003358_s11837-018-2794-3-Figure1-1.png",
        "caption": "Fig. 1. Illustration of (a) vertically (V-) and (b) horizontally (H-) built strips. (c) Long, thin (2 mm) and smooth tensile specimens.",
        "texts": [
            "com. 4.\u2014e-mail: ma.qian@ rmit.edu.au Many novel designs for additive manufacturing (AM) contain thin-walled (\u00a3 3 mm) sections in different orientations. Selective electron beam melting (SEBM) is particularly suited to AM of such thin-walled titanium components because of its high preheating temperature and high vacuum. However, experimental data on SEBM of Ti-6Al-4V thin sections remains scarce because of the difficulty and high cost of producing long, thin and smooth strip tensile specimens (see Fig. 1). In this study, 80 SEBM Ti-6Al-4V strips (180 mm long, 42 mm wide, 3 mm thick) were built both vertically (V-strips) and horizontally (H-strips). Their density, microstructure and tensile properties were investigated. The V-strips showed clearly higher tensile strengths but lower elongation than the H-strips. Hot isostatic pressing (HIP) produced the same lamellar a-b microstructures in terms of the average a-lath thickness in both types of strips. The retained prior-b columnar grain boundaries after HIP showed no measurable influence on the tensile properties, irrespective of their length and orientation, because of the formation of randomly distributed fine a-laths",
            "1,2 Together with its high vacuum, SEBM is therefore well suited to AM of thin-walled (\u00a3 3 mm) titanium components such as turbocharger compressor wheels and knee implants.2\u20134 Although SEBM of Ti-6Al-4V has been well studied, previous research has largely focused on cylindrical and thick plate samples.1\u201312 In fact, only four studies have dealt with SEBM of Ti-6Al-4V thin sections,13\u201316 of which only Ref. 15 (short samples, 50 mm long) and Ref. 16 (the current authors) have recently reported tensile property data to the authors\u2019 knowledge. The reason can be attributed to the challenge of producing long, thin and smooth strip tensile specimens (see Fig. 1), as well as the associated cost from post-processing (polishing and HIP). On the other hand, many novel designs contain thin-walled sections in different orientations and require essential experimental data. This study aims to improve this knowledge gap. Spherical Ti-6Al-4V powder of 45\u2013106 lm in size from Arcam AP&C was used. Table I lists the powder composition and the Al, V and O contents of the as-built strips. Forty strips were built vertically (V-strips) and 40 horizontally (H-strips), as illustrated in Fig. 1a and b, using an Arcam A2 machine by following the Arcam \u2018\u2018Ti6Al4V-Melt50 lm\u2019\u2019 procedure. Each strip measures 180 mm long, 42 mm wide and 3 mm thick. Detailed SEBM parameters including the scanning speed (4500 mm/s) can be found in Ref. 10. Each two adjacent V-strips were spaced by 18 mm. Cellular JOM https://doi.org/10.1007/s11837-018-2794-3 2018 The Minerals, Metals & Materials Society supports (8 mm high, Fig. 1b) were used in building the H-strips. Hot isostatic pressing (HIP) was conducted at 920 C 9 120 min 9 100 MPa in an ASEA Brown Boveri (ABB) QIH-15 unit (hot zone: /200 mm 9 300 mm). The density was measured by the Archimedes method. A GE Phoenix v-tome-x s240 computed tomography (CT) scanner was used to detect any residual pores after HIP. Long tensile samples (Fig. 1c) were machined from as-built strips according to GB/T 228.1-2010 (equivalent to ISO 6892- 1:2009) and tested on an Instron Model 5982 at a strain rate of 10 3 s 1. Three samples were tested for each condition. Polished samples were etched with Kroll\u2019s reagent (2 mL HF, 4 mL HNO3, and 94 mL H2O) and examined using an optical Wang, Tang, Yang, Liu, Jia, and Qian microscope (OM) and a scanning electron microscope (SEM, a Jeol4800, JEOL Ltd. Japan, operated at 7 kV). The as-built strips achieved 4.399 \u00b1 0",
            " The different tensile properties between as-built V- and H-strips (Table III) correspond to their microstructures. The as-built H-strips are featured by lamellar a\u2013b while the as-built V-strips consist of incompletely decomposed a\u00a2-martensite. Consequently, the as-built V-strips are stronger but less ductile than the as-built H-strips. The total energy input for building each V-strip and each H-strip was the same in this study. However, the heat conduction from each H-strip to the substrate occurs through the cellular supports (Fig. 1) while for each V-strip it occurs through the strip itself, which is much more efficient than the former. Therefore the as-built V-strips contained incompletely decomposed a\u00a2-martensite (Fig. 3) while the as-built H-strips had fully lamellar a\u2013b (Fig. 4). After HIP, the resultant a\u2013b lamellae in both the V- and H-strips had the same average a-lath thickness and were both essentially pore-free (Fig. 5). The major difference lies in the length of the retained prior-b columnar GBs. Figure 6 illustrates this sharp contrast"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003356_j.apm.2018.01.018-Figure17-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003356_j.apm.2018.01.018-Figure17-1.png",
        "caption": "Fig. 17. Location of the maximum temperature for the models subjected to over-inflation pressure conditions at 15 km/h: (a) 268 kPa and 22,588 N, (b) 336 kPa and 22,588 N and (c) 336 kPa and 48,656 N.",
        "texts": [],
        "surrounding_texts": [
            "The temperature distribution in the tyre\u2019s cross-section was obtained from each of the operating conditions specified in Table 7 , except for the 450 kPa inflation pressure load conditions. The maximum temperatures in the tyre model for each combination of the operating conditions were of interest. The numerical results were grouped according to the tyre\u2019s state of inflation. The location and value of the maximum temperature in the tyre\u2019s cross-section when it is under-inflated, inflated to the recommended pressure and when it is over-inflated are shown in Figs. 15 , 16 , 17 , respectively. Typically, the state of inflation for earthmover tyres, is a function of the load applied to the tyre, i.e. the recommended inflation pressure for a tyre carrying 9.0 t, is 450 kPa, whereas the recommenced inflation pressure of a tyre carrying 7.4 t, is 336 kPa. The inflation state of the results presented were classified according to the tyre manufacturer recommendations. As expected, the location of the maximum temperature is indicative of the point at which the tyre experiences the most deformation during the specific load condition. Also, the tyre temperature increases as the vertical deflection of the tyre axle relative to the ground and its forward rolling velocity increases. A scatter plot of the tyre\u2019s deflection and the maximum tyre temperature of each of the respective models at 15, 30 and 45 km/h are shown in Fig. 18 . In the mining industry, an operating temperature of below 75 \u00b0C is recommended for safe operation. The recommended operating envelope is indicated on Fig. 18 , and is below 75 \u00b0C. The tyre deflection and rolling velocities that would result in critical temperature conditions are indicated by the red box seen in Fig. 18 . A polynomial fit of the data was used to establish a relation between the maximum tyre temperature and its deflection and velocity conditions. From the results obtained, an estimate of the expected maximum tyre temperature as a function of its velocity, its vertical deflection and the ambient temperature specified in the thermal analysis was established. This function is given as, T max = (21 . 25 V t + 123 . 79)(X disp \u2212 0 . 0132) + T amb (6) where T max is the predicted internal tyre temperature in \u00b0C, V t is the forward velocity of the tyre in km/h, X disp is the vertical tyre deflection in m, and T amb is the ambient temperature of its surroundings in \u00b0C. From Eq. (6) , it is evident that the maximum temperature in the tyre can be controlled by either adjusting the tyre deflection or velocity at the specified ambient conditions. The tyre forward rolling velocity will have the most significant influence on the heat generation in the tyre and it should therefore be managed accordingly."
        ]
    },
    {
        "image_filename": "designv11_13_0002007_1.4029627-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002007_1.4029627-Figure3-1.png",
        "caption": "Fig. 3 Notation for the oscillating wing",
        "texts": [
            " The proposed controller effectively stabilizes the closed-loop system to the constant desired trajectory, about which the system was linearized. To illustrate that the controller can track slowly time-varying trajectories, as well, we present simulation results using the following reference signals: hd1 \u00f0t\u00de \u00bc 0:4 sin t; hd2 \u00f0t\u00de \u00bc 0:3 sin\u00f00:5t\u00de; hd3 \u00f0t\u00de \u00bc 0:5 sin\u00f00:5t\u00de, and hd4 \u00f0t\u00de \u00bc 0:5. The simulation results, with x\u00bc 100 rad/s and zero initial conditions, are presented in Figs. 1 and 2. Consider the 3DOF device depicted in Fig. 3. The device consists of a square plate of mass m and width 2a, with an additional mass m0 that is rigidly fixed at the top of the plate. The configuration is defined relative to a fixed inertial reference frame fi; j; kg. The plate is immersed in a uniform flow moving at speed V1 in the negative y-direction. A linear servo-actuator drives the wing in the direction normal to the flow (the z-direction). The combined motion of the plate in the z and h directions, which we call 071004-6 / Vol. 137, JULY 2015 Transactions of the ASME Downloaded From: http://dynamicsystems",
            "org/about-asme/terms-of-use \u201cflapping,\u201d creates aerodynamic forces which can be used to control the motion of the system in the y direction. The aerodynamic forces are determined using a quasi-steady aerodynamic model. Since the motion of the plate causes acceleration of the fluid in the direction normal to the plate, a parameter c is included to account for the added mass due to the fluid. Finally, a linear torsional spring (with stiffness kt) and a linear damper (with damping coefficient bt) resist the plate\u2019s angular motion, as indicated in Fig. 3. Using Lagrange\u2019s method [18, Chap. 2], one finds that the dynamic model takes the form [19] M\u00f0h\u00de \u20acz \u20ach \u20acy 0 @ 1 A \u00bc fz fh fy 0 @ 1 A |fflfflffl{zfflfflffl} f \u00fe uz\u00f0t\u00de 0 0 0 @ 1 A |fflfflfflfflffl{zfflfflfflfflffl} u\u00f0t\u00de (33) where the inertia matrix M\u00f0h\u00de \u00bc m0 \u00fe m\u00f01\u00fe c cos2 h\u00de ma\u00f01\u00fe c\u00de cos h 1 2 cm sin 2h ma\u00f01\u00fe c\u00de cos h I0 \u00fe Ip \u00fe ma2\u00f01\u00fe c\u00de ma\u00f01\u00fe c\u00de sin h 1 2 cm sin 2h ma\u00f01\u00fe c\u00de sin h m0 \u00fe m\u00f01\u00fe c sin2 h\u00de 0 B@ 1 CA and where f \u00bc faeroz faeroh faeroy 0 B@ 1 CA\u00fe cm _h\u00f0 _y cos 2h _z sin 2h\u00de \u00fe ma\u00f01\u00fe c\u00de _h2 sin h cm _y _z cos 2h\u00fe 1 2 cm\u00f0 _y2 _z2\u00de sin 2h bt _h kth mga sin h cm _h\u00f0 _y sin 2h\u00fe _z cos 2h\u00de ma\u00f01\u00fe c\u00de _h2 cos h \u00f0m\u00fe m0\u00deg 0 BB@ 1 CCA with faeroz ; faeroh , and faeroy representing the aerodynamic forces and moment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003607_j.procir.2018.08.113-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003607_j.procir.2018.08.113-Figure2-1.png",
        "caption": "Fig. 2. Tool setup for the shear test with AE in the pressure zone",
        "texts": [
            " Furthermore, the influence of different process routes on the shear bonding strength with an AE at the tension zone is investigated in [5]. Based on these results the influence of different stress states at the AE during the warm bending process on the shear bonding strength is unclear. Therefore, the comparison of different stress states is part of this work. Compared to testing hybrid parts with AE in the tension zone [5], the tool is adapted for characterizing hybrid parts with AE in the pressure zone. The tool used in this work is shown with relevant components in Fig. 2. A prism and a clamping prism representing the shape of the formed sheet metal part are developed to fix the specimen during the test. Furthermore, the shape of the punch is modified to shear the AE in the concave geometry. The punch is guided by a plate, which is fixed on the clamping plate. A hard stop at the bottom of the prism prevents slipping of the specimen. The tool is implemented in a universal testing machine walter+bai FS-300 with a maximum force of 300 kN. The shear tests are conducted with a punch velocity of 5 mm/min"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002019_0954406214525364-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002019_0954406214525364-Figure4-1.png",
        "caption": "Figure 4. Schematic of the finger seal leakage clearance.",
        "texts": [
            " Dynamic performance of the finger seal with a stack of finger elements An approximate annular leakage clearance is shaped between the finger element and the rotor due to frictions between the aft cover plate and finger element and between finger elements during running Ff i\u00fe1\u00f0 \u00de \u00bc Fs i\u00fe1\u00f0 \u00desgn Fi\u00fe1\u00f0 \u00de Fi\u00fe1 Fd i\u00fe1\u00f0 \u00desgn _xi\u00fe1 t\u00f0 \u00de\u00f0 \u00de 8>< >: Fi\u00fe1 5Fs i\u00fe1\u00f0 \u00de, _xi\u00fe1 t\u00f0 \u00de 4\" Fi\u00fe1 5Fs i\u00fe1\u00f0 \u00de, _xi\u00fe1 t\u00f0 \u00de 4\" other 14i4n 1\u00f0 \u00de \u00f09\u00de at WEST VIRGINA UNIV on April 11, 2015pic.sagepub.comDownloaded from condition. The schematic diagram of the finger seal leakage clearance considering interactions between finger elements is shown in Figure 4. For finger seal, one cyclically symmetric structure (defined in the Equivalent dynamic model section) is investigated considering the cyclic symmetry feature of the seal configuration. During one rotary cycle of the rotor, the difference in displacement amplitudes between the rotor and the cyclically symmetric structure is the instantaneous leakage clearance. Thus, air flow mass through the leakage clearance is the instantaneous leakage mass of cyclically symmetric structure. So the total leakage flow mass of finger seal is obtained by adding up the leakage flow mass of each cyclically symmetric structure during one operating cycle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003620_j.matpr.2018.06.241-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003620_j.matpr.2018.06.241-Figure3-1.png",
        "caption": "Fig. 3 Bicycle subjected to load conditions (i) and (ii)",
        "texts": [
            "/ Materials Today: Proceedings 5 (2018) 18920\u201318926 Devaiah B.B et al./ Materials Today: Proceedings 5 (2018) 18920\u201318926 18923 (i)Static Start up- the rider is on the bicycle applying a load of 700N about to start pedalling but at rest. Aerodynamic, rolling and gyroscopic forces are assumed to be negligible. The bicycle is in vertical equilibrium. (ii)Steady State pedalling- the cyclist is seated on the bicycle and applying a force of 200N due to leg dynamics. The load is assumed to be concentrated at the bearing as shown in above Fig 3 (i -ii). (iii)Vertical Impact- vertical impact loads are represented by multiplying the cyclist\u2019s weight by some amount of G factor. In this case a factor of 2G is taken taking the load to 1400N which is the necessary case when an object falls from an infinitesimal height onto a rigid surface as shown in below Fig 4 and Fig 5. (iv)The case that is presented now is from [2]. Here the loads are simulated for the load bumps occurring at the front wheel. A resultant load of 2700N is transmitted at the rigid links which are then connected to the axle and then to the frame via the fork"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure7.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure7.1-1.png",
        "caption": "Fig. 7.1 Fragment of the cylindrical conducting layer",
        "texts": [
            " Thus, in order to construct an electric machine equivalent circuit corresponding to under-load operation, the magnetic circuit regions (layers) characterized by the presence of eddy or external currents should be replaced by equivalent circuits taking into account these currents. Therefore, in this chapter, equivalent circuits of passive and active conducting layers are considered which represent the \u201cbuilding blocks\u201d of an electric machine magnetic circuit model. In this chapter we use the provisions described in [1\u201316]. For purposes of clarity, we consider the passive conducting layer as the wound part of the rotor tooth region, the conditional design scheme of which is shown in Fig. 7.1. It is assumed that this rotor region has been provided with a squirrel-cage type winding. As was shown in Chap. 6, the layer equations have the form \u00a9 Springer International Publishing Switzerland 2015 V. Asanbayev, Alternating Current Multi-Circuit Electric Machines, DOI 10.1007/978-3-319-10109-5_7 227 In (7.1), the values of EzR1,HzR1 and EzR2,HzR2 represent the electric and magnetic field strengths on the outer and inner surfaces of the considered layer (Fig. 7.1). In the case of eddy currents present in the rotor winding, the constants of the system of equations (7.1) represent complex numbers [4, 12\u201317]. It is assumed that the constants z11, z21, z22 and z12 in (7.1) are symmetrical, i.\u0435., z11\u00bc z22 and z12\u00bc z21. As was shown in Chap. 5, this condition is provided when the calculated values of HzR1 and HzR2 determined by (5.109) and (5.111) are used in system of equations (7.1). System of equations (7.1) can be used to analyze the processes taking place in the conducting layer",
            "1 Passive Layer Circuit Loops: Methods for Obtaining 229 From (7.2), we can obtain EzR1 \u00bc Z\u03c4zR1HzR1 \u00fe EzR0 EzR2 \u00bc Z\u03c4zR2HzR2 \u00fe EzR0 \u00f07:4\u00de Equations (7.4) describe the conducting layer circuit loops determined in relation to the value of emf EzR0. System (7.4) represents the coupling equations demonstrating the mutual connection of the magnitude of emf EzR0 with the values of emf EzR1 and EzR2 determined on the outer and inner layer surfaces, respectively. On the basis of system of equations (7.2) obtained for the layer shown in Fig. 7.1, the \u0422-circuit given in Fig. 7.3 arises. It follows from this T-circuit that the value of emf EzR0 depends on the voltage drop across the impedance Z\u03c4zR1, and it is determined from (7.4) as EzR0\u00bcEzR1 Z\u03c4zR1HzR1. Generally, in this connection, the use of equation EzR0\u00bc ZzR0HzR0 for an analysis of the processes caused by the own current HzR0 in the considered layer is associated with certain disadvantages. So, for example, under the condition of strong skin effect when the voltage drop across impedance Z\u03c4zR1 increases sharply, equation EzR0\u00bc ZzR0HzR0 is no longer valid",
            " In accordance with this principle, the total layer field can be represented as the superposition of two fields: the field created by the own layer current and the field caused by the total current flowing in the layers located below the considered layer [4]. The processes in the layer can be considered in relation to the surface through which the electromagnetic energy penetrates the layer (for the rotor layer outer or upper surface). In order to implement the field decomposition principle, we assume that the magnitude of current HzR1 located on the outer layer surface (Fig. 7.1) can be represented as the geometric sum of the own layer current and the current proportional to the total current flowing in the layers located below the considered layer. Therefore, current HzR1 can be represented as a result of the superposition of EzR2 0 0 HzR0 EzR0 HzR1 HzR2 EzR1 Z\u03c4zR1 Z\u03c4zR2 ZzR0 \u00b7 \u00b7 Fig. 7.3 T-circuit representation of the conducting layer 230 7 Passive and Active Conducting Layers: The Circuit Loops two currents (when EzR1\u00bc const): current HzR and current ( H zR2), i.\u0435., HzR1\u00bcHzR + ( H zR2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002445_ijmmme.2016010103-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002445_ijmmme.2016010103-Figure4-1.png",
        "caption": "Figure 4. Dynamic structural analysis of transmission gearbox",
        "texts": [
            "5 evaluate the results of structural and steady state thermal analysis. Structural results evaluate the performance of transmission gear train on strength point. Inertia and damping effects was not considered for simulation. When vehicle was running at full loading and 1500 rpm and generates 245 Nm torque. Torque and rotational velocity were applied for full dynamic loading. At maximum torque condition the gear meshing causes heat generation in large value which was studied in this research work. Von-mises stress and strain were mentioned in figure 4. Figure 4 (a) the simulation results shows that the equivalent (Von-Mises) stress distribution is within safe limit at its minimum value. The shear elastic strain due to dynamic load shows deformation. This deformation is safe (figure 4 (b)). No high deformation regions were identified. Figure 4 structural results show that the gear train design is safe to sustain the applied dynamic load. Transient structural analysis was performed at 1500 rpm when vehicle was on high gearing (third gear selection) and producing 119.31 Nm torque. The time period of full loading was 9 seconds and dynamic behaviour of 4-speed transmission gear box was evaluated. Figure 4 (c and d) shows transient results. Figure 4 (c) explains the total deformation variation in gears. When gears are in meshing at full loading the red hues region shows stress concentration where deformation is maximum. The blue hues signify the minimum deformation level and light green hues shows the regions where deformation is average. Maximum value of total deformation (red hues) is 0.126 mm within limit. Figure 4 (d) shows maximum principal stress variation on 4-speed transmission gear box. Maximum principal stress available on gears is 5.346e8 Pa. Blue hues show the principal stress variation, which is uniform in nature. A uniform stress signifies that failure due stress concentration will be unavailable. The present research work concerns with the transmission assembly of medium duty trucks. In earlier study the authors have considered only gear tooth or simple geometry of transmission but here we have simulated the full assembly to highlight the effects of different operating conditions (load, rotational speed, lubrication)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003234_1350650117738395-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003234_1350650117738395-Figure2-1.png",
        "caption": "Figure 2. Example of temperature increase in a molybdenum coated synchronizer.23",
        "texts": [
            "19 measured and simulated the focal surface temperature of a brass synchronizer, and found that above 180 C the risk for clashing as well as the wear rate increased. Both Spreckels20 and Neudo\u0308rfer21 compared focal surface temperature finite element (FE) simulation with test for molybdenum coated synchronizer, and for sinter bronze and carbon lined synchronizers, respectively. These three investigations suggest that the focal surface temperature is a critical parameter for synchronizers. A FE based thermomechanical simulation model has previously been developed5 and validated.22 An example of the temperature increase is available in Figure 2. The model can be used to study the contact temperature of nominally smooth synchronizer surfaces. Since the surfaces are considered nominally smooth, no asperity contact and hence no flash temperatures are included. In addition to the FE model, a one-dimensional (1D) thermal model of a synchronizer has been developed to determine the average surface temperature during synchronization.23 The only geometric dimension considered in the 1D thermal model is the depth below the surface, therefore no gradient in the axial direction is available"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000203_j.ijmecsci.2013.11.010-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000203_j.ijmecsci.2013.11.010-Figure8-1.png",
        "caption": "Fig. 8. Relative sliding between tube and thread.",
        "texts": [
            " (24) can be derived from the following equation: \u222c\u03a9uP\u2032 dx dy\u00bc\u222c\u03a9 Nx \u2202u \u2202x 2 \u00feNy \u2202u \u2202y 2 \" # dx dy; \u00f026\u00de where the region \u03a9\u00bc f\u00f0x; y\u00dejjxj=lx\u00fejyj=lyr1g, and the forces Nx and Ny per unit length can be written in terms of sx;rub and s\u03b8;rub as follows: Nx \u00bc sx;rubtR; N\u03b8 \u00bc s\u03b8;rubtR: \u00f027\u00de Substituting Eq. (24) into Eq. (26), the parameter C can be written by the following: C \u00bc 32P\u2032 \u03c04 Nx l2x \u00feNy l2y \" #: \u00f028\u00de Firstly, the magnitude of relative sliding between the deformed tube and the surrounding threads is discussed. Fig. 8 shows the movement of a thread from the position AD to the position A\u2032D\u2032 as the angle of rotation changes from \u03b3 to \u03b3\u00fed\u03b3. Consider an arbitrary point M along the centerline of the thread and another point N separated from point M by a distance s as shown in Fig. 8. For the position AD of the thread, the coordinates of the point N can be written as follows: xN \u00bc xM\u00fes sin \u03b3; yN \u00bc yM\u00fes cos \u03b3; \u00f029\u00de where \u00f0xM ; yM\u00de are the coordinates of the point M . When the point N is attached completely to the thread, and the thread deforms from the angle \u03b3 to \u03b3\u00fed\u03b3, the coordinate values termed xNt and yNt can be expressed as follows: xNt \u00bc k1xM\u00fek2s sin \u03b3; yNt \u00bc k2yM\u00fek1s cos \u03b3; \u00f030\u00de where parameters k1 and k2 are given by the following: k1 \u00bc jOD\u2032j jODj ffi1 sin \u03b3 cos \u03b3 d\u03b3; k2 \u00bc jOA\u2032j jOAj ffi1\u00fe cos \u03b3 sin \u03b3 d\u03b3: \u00f031\u00de On the other hand, when the point N is attached completely to the rubber, the coordinate values at the point N after deformation are denoted by \u00f0xNr ; yNr\u00de, and these values can be expressed as follows: xNr \u00bc k1\u00f0xM\u00fes sin \u03b3\u00de; yNr \u00bc k2\u00f0yM\u00fes cos \u03b3\u00de: \u00f032\u00de Therefore the magnitude of the relative sliding denoted by dutr can be written as follows: dutr \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0xNr xNt\u00de2\u00fe\u00f0yNr yNt\u00de2 q \u00bc 2s j cos 2\u03b3j sin 2\u03b3 d\u03b3: \u00f033\u00de It can be understood from Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001183_tmag.2017.2661381-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001183_tmag.2017.2661381-Figure3-1.png",
        "caption": "Fig. 3. 3-D mesh of the 37 kW motor\u2019s axial section.",
        "texts": [
            " As expected, the powers predicted with the ill-posed system in (2) is highly sensitive to the measurement error and differ widely from the actual values. On the other hand with constrained LS, the results match the actual power distribution very well. The lumped network is calibrated against the temperatures measured on the motor\u2019s frame, stator endcoils, and stator winding temperature at full-load. These are used to tune the thermal network [8], and the resulting calibrated analytical thermal coefficients are applied to the 3-D FE model of the motor built in COMSOL seen in Fig. 3. This explicit definition of the boundary heat fluxes makes the heat-transfer computation in coupled fluid-solid interfaces easier and faster, and results in a complete steady-state thermal profile of the motor. The idea is to use mean domain temperatures from these numerical results in place of the actual measurements to further assess the effectiveness of the constrained LS method in inverse mapping. The new temperature vector now has 18 such average values corresponding to the unique machine parts depicted in the thermal network"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001034_jfm.2015.661-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001034_jfm.2015.661-Figure2-1.png",
        "caption": "FIGURE 2. (a) Geometric definition of an elliptical foil with AR = b/c = 0.5; (b) representative grid distribution around the foil.",
        "texts": [
            " Asymmetry still arises at smaller values of \u03b2 for given KC, but the asymmetry is now of S-type for all values of KC, with a synchronous signal in the horizontal force on the foil at twice the frequency of the primary oscillation. We discuss the implications of these results for the transition to flying locomotion, and briefly draw our conclusions in \u00a7 5. 2. Problem description and numerical method 2.1. Problem description We consider elliptical foils with major axis c and minor axis b, such that AR=b/c61, with uniform mass density \u03c1s, as shown in figure 2(a). The elliptical foil translates in the infinite x\u2013y plane through a two-dimensional viscous fluid of density \u03c1 and kinematic viscosity \u03bd. Although variations in the density ratio \u03c1s/\u03c1 undoubtedly affect the flow dynamics after symmetry breaking (see e.g. Alben & Shelley 2005), here we are exclusively interested in the initial behaviour very close to the transition boundary, and so for simplicity, we keep the density ratio fixed at the single value \u03c1s/\u03c1 = 10. As noted above, we impose a vertical oscillation of the centre of the foil so that ys(t)=A sin(2\u03c0f0t)",
            " We have found that requiring these two conditions yields time accurate and robust results. To validate the spatial resolution we use, we have carried out a grid-independence study on a purely oscillating elliptical foil with aspect ratio AR= 1.0, i.e. a circular cylinder in two-dimensional space, analogously to the study discussed in detail in Deng, Caulfield & Shao (2014). We find that meshes with approximately 50 000 cells provide satisfactory and consistent accuracy in space. As an example, we plot the grid near an elliptical foil with AR = 0.5 in figure 2, which shows the gradual increase of the mesh size from the foil boundary. The domain is defined as a circle with a radius 20c. Pressure and all components of the velocity gradient tensor are set to zero on the boundary of the domain. Further confidence in the fidelity of our simulations is gained by the good agreement with previous numerical simulations by Elston et al. (2001) of the calculated transition boundary for flow around an AR= 1.0 foil shown in figure 1, particularly for low \u03b2 numbers",
            " Instability occurs when a multiplier leaves the unit circle, |\u00b5| > 1, or equivalently when the real part of a Floquet exponent becomes positive. The technique we use for Floquet stability analysis is a Krylov subspace method that examines the stability of the linearized Poincar\u00e9 map for the perturbation flow, and is detailed in Elston et al. (2004, 2006) who, as already mentioned, considered the Floquet stability of the flow around a circle with AR= 1. The reflection symmetry of the base flow is enforced by solving in a half domain (see figure 2), with symmetry boundary conditions along the x= 0 boundary. The base flow is integrated in time for 30 cycles, when it reaches a periodic state. It is then projected onto the full domain, and stored for Fourier time interpolation. We store 64 time slices, equi-spaced in time over the base flow period T , for reconstruction of the base flow. It should be noted that for two-dimensional Floquet analysis in the current problem, there is difficulty resolving stable modes, |\u00b5|< 1, while unstable modes are resolved without difficulty, and the location of marginal stability can be estimated by extrapolation to |\u00b5| = 1 (Elston et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000616_aim.2012.6265962-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000616_aim.2012.6265962-Figure5-1.png",
        "caption": "Fig. 5. State and gait\u2019s transition",
        "texts": [
            " (16) when supporting-foot begins slipping; and jumping motion by adding further variable of upright direction z0 to q when jumping represented by Fig. 4. Table II indicates all possible walking gaits regarding contacting situations\u2014surface-contacting (S), point-contacting (P) and Floating (F)\u2014of supporting-foot (S.F.) and floatingfoot or contacting-foot (F.F.). The Table is basically divided into three blocks representing the gait\u2019s varieties from a point of states of supporting-foot, such as, \u201cStop,\u201d \u201cSlip\u201d and \u201cAir.\u201d Figure 5 depicts all possible gait transition of bipedal walking based on event-driven, which indicate that appropriate dynamics and variables are selected and applied according to the phase or state, which are listed in Table II. In the state that has ramification such as state (II) in Fig. 5 into state (II\u2032) or (III), the gait is switched to next state in case that auxiliary switching condition written above the allow in the figure indicating phase transient is satisfied. In the gait transition from (III) or (III\u2032) to (IV), supporting-foot is switched from one foot to the other foot with renumbering of link, joint and angle\u2019s number. What the authors want to emphasize here is that the varieties of this transition completely depend on the solution of dynamics shown as Eq. (10) or Eq. (16). A condition that heel of supporting-foot detaches from the ground in Fig. 5 (I), (II), (III) to (I\u2032), (II\u2032), (III\u2032) is discussed. For this judging, 2f2 and 2n2 calculated from Eqs. (5), (6) in case of i = 2 are used. Firstly, coordinates of 2f2 and 2n2 represented by Fig. 6 (a) are converted from \u03a32 to \u03a3W . Then, projection to z-axis of the force and projection to x-axis of the torque are derived by using unit vector ex = [1, 0, 0]T and ez = [0, 0, 1]T as: Wf2z = eT z (WR2 2f2), Wn2x = eT x (WR2 2n2) like Fig. 6 (b). Given that supporting-foot\u2019s contacting points are to be two of toe and heel as shown Fig. 6 (c), when forces acting on the toe and heel are defined as ff , fr, these forces must satisfy the following equations. Wf2z = ff + fr (17) Wn2x = \u2212ff \u00b7 lf + fr \u00b7 lr (18) We can calculate ff and fr as Eq. (19) and supporting-foot begins to rotate around the toe like Fig. 6 (d) when value of fr becomes negative. ff, r = lr \u00b7 Wf2z lf + lr \u00b1 Wn2x lf + lr (19) When floating-foot attaches to ground, we need to consider bumping motion [16]. Figure 5 has two kinds of bumping concerning heel and toe. We denote dynamics of bumping between the heel and the ground below. By integrating Eq. (12) under \u03c4n = 0 in time, equation of striking moment can be obtained as follows. M(q)q\u0307(t+1 ) = M(q)q\u0307(t\u22121 ) + (jT c \u2212 jT t K)Fim (20) Eq. (20) describes the bumping in z-axis of \u03a3W between the heel and the ground. q\u0307(t+1 ) and q\u0307(t\u22121 ) are angular velocity after and before the strike respectively. Fim = limt\u22121 \u2192t+1 \u222b t+1 t\u22121 fndt means impulse of bumping. Motion of the robot is constrained by the followed equation that is given by differentiating C1 by time after the strike",
            " Here, the force exerted on the head to minimize \u03b4\u03c8(t) = \u03c8d(t) \u2212 \u03c8(t) calculated from HT Hd \u2014the pose deviation of the robot\u2019s head caused by gravity force and walking dynamical influences\u2014is considered to be directly proportional to \u03b4\u03c8(t). The joint torque \u03c4h(t) that pulls the robot\u2019s head up is given the following equation: \u03c4h(t) = Jh(q)T Kp\u03b4\u03c8(t), (24) where Jh is Jacobian matrix of the head pose against joint angles and Kp means proportional gain similar to impedance control. We use this input to compensate the falling motions caused by gravity or dangerous slipping motion happened unpredictably during all walking states in Fig. 5. Notice that the input torque for non-holonomic joint like q1 (foot tip joint), \u03c4h1 in \u03c4h(t) in Eq. (24) is to be set as zero since it is free joint. Under the environment that sampling time was set as 3.0 \u00d7 10\u22123 [sec] and friction force between foot and the ground as ft = 0.7fn, the following simulations were conducted. In regard to simulation environment, we used \u201cBorland C++ Builder Professional Ver. 5.0\u201d to make simulation program and \u201cOpenGL Ver. 1.5.0\u201d to display humanoid\u2019s time-transient configurations",
            " However, amplitude of arms\u2019s swing became large spontaneously and converged to a certain amplitude and period. Therefore, we can say that both arms\u2019 swing were caused by interactions of walking dynamics. Figure 11 shows state transition generated by the humanoid\u2019s dynamics, both feet\u2019s position in y-z plain and displacement of ZMP during one walking step. In this simulation, the humanoid walked in accordance with the following path: (I) \u2192 (I\u2032) \u2192 (II\u2032) \u2192 (III\u2032) \u2192 (IV) \u2192 (I) \u2192 \u00b7 \u00b7 \u00b7 . This transition was selected among all possible transient in Fig. 5 by the solution of dynamics represented by Eqs. (10), (16), initial condition and input torque. That is, the path of transition will be changed easily by these factors. Moreover, Fig. 11 denotes that ZMP moves forward and reaches the edge of supporting-foot while the other foot in the air, meaning that the robot is tipping over, which does not appear in ZMP-based walking. We think that this kind of natural walking is caused by the effect of visual feedback as shown in the following subsection. We assume that two patterns of supporting-foot\u2019s contacting and input torques based on Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001985_ecce.2015.7309702-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001985_ecce.2015.7309702-Figure7-1.png",
        "caption": "Fig. 7. Oil inlet/outlet, and lubrication oil control valve of custom-built small scale journal bearing test setup for 380 V, 5.5 kW, 2 pole induction motor (motor B)",
        "texts": [
            " Holes were drilled in the axial and radial direction in the upper part of the bearings to guide the incoming oil from the adapter to the shaft-bearing clearance, as shown in Fig. 6(a). The path of the oil flow through the inlet, channel, adapter, bearing, shaft, channel, and outlet is shown in Fig 5 and in the bearing design drawing shown in Fig. 6(b). It is also critical to control the \u201cflow\u201d of oil to maintain the oil film in the bearing for operation of the motor within its thermal limits. The flow of ISO viscosity grade 68 oil was controlled by feeding the oil into the grease inlet of the end shield through a valve, as shown in Fig. 7, and the oil leaving the grease outlet was circulated back to the inlet. Rubber seals were placed on the shaft to minimize the leakage of oil. The motor was fixed to a 42.0 kg steel baseplate to eliminate other sources of vibration due to structural looseness of the foundation, as shown in Fig. 7. IV. EXPERIMENTAL RESULTS Accelerometer measurements were obtained at a sampling frequency of 800 Hz from two identical sealless pump motor units described in III.A operating at 70% rated load at 3495 rpm (motors A1-A2). The results of vibration spectrum analysis for a frequency range of 0 to 400 Hz for motors A1 and A2 are shown in Fig. 8(a)-(b), respectively. The overall vibration level was higher for motor A2 at 0.068 in/sec when compared to that of motor A1 at 0.053 in/sec. The two largest vibration components for motor A1 were the rotor speed frequency given by (1) at fr,vib=58",
            " From the test results and visual inspection, it can be concluded that the clearance of motor A2 was increased due to shaft and bearing contact. It is suspected that the excessive clearance caused oil whirl instability at the time of vibration and current measurements, and intermittent contact between the journal and bearing. Visual inspection shows that shaft/bearing rub can lead to catastrophic motor and pump system failure, if left undetected. A 500 mV/g commercial accelerometer was installed on the DE of motor B for measuring vertical acceleration to perform vibration analysis, as shown in Fig. 7. Vibration and current spectrum analysis were performed on motors B1-B3 with small scale journal bearings described in III.B with bearing clearances of 60, 75, and 90 \u03bcm, to observe if MCSA can be used to detect increased clearance in the journal bearing. The vertical vibration and stator current data were simultaneously measured at a sampling frequency of 6.4 kHz when the motor was operated under no load condition (fr,vib=59.94 Hz, 3596.4 rpm). The components related to the increased clearance at 1/2 and 1/3 of the rotor speed (19"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001192_s10010-017-0250-0-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001192_s10010-017-0250-0-Figure2-1.png",
        "caption": "Fig. 2 Generating conical surface in a",
        "texts": [
            " b Position relation among coordinate systems O1 1 ko1=k1 jo1 kd S=1 S=2 Lw/2 Lw/2 Enveloping spiroid workpiece Grinding wheel i flank e flank a b io1 jo1 ko1=k1 1 ra O1 Ood jod jo1 jd iod = id kd kod ia jd ja id Oaka ia id ja jd ka Oa Od i1 j1 r1 rd 1 \u03b3 m ad p\u03d5 Oaka ja kd Od id Oa ka ia g1 g2 rg ( 2) P Od ko1 io1 kod iod Od jod kod jd kd \u03b3 m \u03b5 1(\u03b5 2) \u03b4 \u03b4 \u03b4 \u03c6 \u03c9 \u03b5 \u03b5 Moreover, the theory and technique proposed can easily be applied to other gearings with complex geometry. A coordinate system a n OaI\u2212!i a; \u2212! j a; \u2212! ka o is fixed with a disk-shaped conical grinding wheel, and the unit vector \u2212! k a is along its axial line as depicted in Fig. 2. By means of the sphere vector function [13], the equation of the generating conical surface, P g, can be represented in a as: ! r a a = u ! ma ; \u0131g = usin\u0131gcos ! i a + usin\u0131gsin ! j a + ucos\u0131g ! ka; u > 0,0 < 2 ; (1) where u and are the two parameters of P g. Here the symbol \u0131g denotes the half taper angle of the conical grinding wheel. According to differential geometry, the unit normal vector of the generating conical surface, P g, can be worked out from Eq. 1. The result is: K ! n a = ! n a ; \u0131g = cos\u0131gcos "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000683_icra.2013.6631258-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000683_icra.2013.6631258-Figure1-1.png",
        "caption": "Fig. 1. Applications of dynamic reconfiguration manipulability for (a) redundant manipulator and (b) humanoid robot.",
        "texts": [
            "jp humanoid robot in this paper, proposing dynamical reconfiguration manipulability (DRM) concept, which is a measure of how much a dynamical system can potentially produce a motion in a workspace with normalized input torque, by combining the dynamic manipulability [12] with reconfiguration manipulability [13]. This new measure represents how much the dynamical system of robots possesses shapereconfiguration acceleration ability in workspace by unit torque input for all joints during executing primary tasks. The DRM have been applied to a humanoid robot, whose prior task be allocated to sustain head position or posture to be close to desired it as much as possible. The concept is shown in Fig. 1(b) and the dynamic reconfiguration manipulability ellipsoid (DRME) of floor-fixed four link robot is shown in Fig. 1(a), and we call ellipsoid as including ellipse and line segment in this paper. Simulations show the DRME varies with the human\u2019s waist position, suggesting that strategies can be applicable such as the whole configuration control can be modified through DRME based on requirement of secondary task other than walking. Walking on uneven ground is shown in Fig. 2. Considering 978-1-4673-5643-5/13/$31.00 \u00a92013 IEEE 4779 condition of task-1 (maintaining head-height) and task-2 (lowering waist-height) is given"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002641_j.ijepes.2016.04.050-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002641_j.ijepes.2016.04.050-Figure3-1.png",
        "caption": "Fig. 3. The reference frames used for the BDFRG model under balanced condition.",
        "texts": [
            " The BDFRG mathematical model under balanced condition The space vector model for the BDFRG is as follows [5,41]: vp \u00bc Rpip \u00fe dkp dt \u00fe jxpkp \u00f01\u00de vs \u00bc Rsis \u00fe dks dt \u00fe j\u00f0xr xp\u00deks \u00bc Rsis \u00fe dks dt \u00fe jxsks \u00f02\u00de kp \u00bc Lpip \u00fe Lpsi s \u00f03\u00de ks \u00bc Lsis \u00fe Lpsi p \u00f04\u00de where Rp, Rs, Lp, Ls and Lps are respectively primary resistance, secondary resistance, primary inductance, secondary inductance and primary to secondary mutual inductance. xp and xs are the primary and secondary space vectors in the reference frames rotating at xp andxs, respectively. xp, xs and xr are the angular frequency of the primary, the angular frequency of the secondary and electrical angular velocity of the rotor, respectively, where the fundamental angular relation between them can be expressed as [2]: xr \u00bc xp \u00fexs \u00f05\u00de The reference frames which are used in the model are shown in Fig. 3 [5]. Finally, the electrical torque expression for the BDFRG is as follows [41]: Te \u00bc j 3 4 Pr kpi p k pip n o \u00f06\u00de The BDFRG mathematical model under unbalanced grid voltage condition In this section, the mathematical model of the BDFRG under unbalanced grid voltage condition has been derived. According to the symmetrical component theory, every unbalanced threephase system can be divided into three balanced three-phase systems: (a) positive sequence, (b) negative sequence, and (c) zero sequence",
            " Referring (11) to the secondary reference frame (by multiplying at e jxst) yields: xs \u00bc xs1ejhxs\u00fe \u00fe xs2ejhxs ej2xpt \u00f012\u00de Eq. (12) can also be rewritten as: xs \u00bc x\u00fes \u00fe x s e j2xpt \u00f013\u00de where x\u00fes and x s are the positive and negative sequence components of xs in the respective positive and negative reference frames, which are rotating at xr xp =xs and xr +xp, respectively. The reference frames which are proposed to model the BDFRG under unbalanced condition are shown in Fig. 6. The dq, dqp and dqs frames in Fig. 6 are the same were in Fig. 3. xp and xs have not been shown again in order to prevent intricacy. Fig. 6 also shows the dq\u00fe p and dq\u00fe s frames, which are respectively overwritten on the dqp and dqs frames, and the dq p and dq s frames, which are respectively rotating at xs and xr \u00fexp. Now, with substituting (12) and (13) into (1)\u2013(4), the space vector model for the BDFRG in the positive and negative sequences can be obtained. The BDFRG model in the positive sequence is as follows: v\u00fe p \u00bc Rpi \u00fe p \u00fe dk\u00fep dt \u00fe jxpk \u00fe p \u00f014\u00de v\u00fe s \u00bc Rsi \u00fe s \u00fe dk\u00fes dt \u00fe j\u00f0xr xp\u00dek\u00fes \u00bc Rsi \u00fe s \u00fe dk\u00fes dt \u00fe jxsk \u00fe s \u00f015\u00de k\u00fep \u00bc Lpi \u00fe p \u00fe Lpsi \u00fe s \u00f016\u00de k\u00fes \u00bc Lsi \u00fe s \u00fe Lpsi \u00fe p \u00f017\u00de It can be seen that the space vector model in the positive sequence under unbalanced condition (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure14-1.png",
        "caption": "Figure 14: Reduced strut uparching and stress peaks in flux of force optimized lattice structure with straightened struts",
        "texts": [
            " Together with the applied force of 300 N, a stiffness of 2317 N/mm results for the structure. This means an improvement of 7 % compared to the curved structure and 9 % compared to the periodic geometry. This leads to a value of 27.7 N/(mm*g) for the stiffness to mass ratio. Hence, an additional increase of 7 % results in contrast to the curved structure. Maximum Stress. Way more important than the enhancement of the stiffness is a decrease of the maximum of the appearing von-Mises stress compared to the curved structure. Figure 14 shows the sectional view at cutting plane 1. T In order to make the results more comparable, the colour scale has been chosen similar to figure 11. It is recognizable, that no or only marginal uparching appears in the single struts. This leads to significantly reduced stress states and stress peaks. The result is a maximum von-Mises stress of 31.9 N/mm\u00b2. Under the assumption, that this maximum stress linearly depends on the applied force, a maximum load of 516.8 N can be calculated for the obtaining of the limit of elasticity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001136_s00170-017-0625-2-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001136_s00170-017-0625-2-Figure1-1.png",
        "caption": "Fig. 1 Overall structure of the hybrid machining device that includes the laser deposition function marked by a dashed line [34]",
        "texts": [
            " In this work, the morphology and evolution of the microstructure in SUS304 austenite stainless steel depositions are further studied. According to the structures that commonly exist in additive manufacturing components, the thin wall and ladder block structures were deposited as two types of samples in this investigation. All the deposited samples are observed in two orthogonal sections. A hybrid machining device based on laser deposition and milling was employed in this study. The overall structure of the device is shown in Fig. 1 and the part involved in laser deposition is marked with dashed lines. The laser beam is produced by a Nd:YAG laser generator of 1.06-\u03bcm wavelength. The deposition process is under the protection environment of Ar gas. The metal substrate and wire used in this study were both composed of SUS304 stainless steel. The diameter of the metal wire was 0.6 mm. The chemical composition of SUS304 is provided in Table 1. The substrate was an industrial cold rolling plate. Before the experiment, the surface of the substrate was polished and cleaned with acetone and alcohol to remove impurities such as oil stains and surface oxide"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003387_s10586-018-2292-y-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003387_s10586-018-2292-y-Figure4-1.png",
        "caption": "Fig. 4 Full vehicle suspension system",
        "texts": [
            " The car body moves along with the movement of the hydraulic cylinder body, thus realizing leveling [18]. For active suspension control, the l/4 or 1/2 vehicle model is usually adopted. But considering the time lag between the front and rear wheel input, as well as the relationship between the coherence of wheel rut, the whole vehicle mathematical model with authenticity can be established to meet the needs of people in practical applications. For the whole vehicle suspension system, a nonlinear dynamic model is established as shown in Fig. 4. The body is treated as a rigid body, and the tire is reduced to a spring with equivalent stiffness. The 4 wheels and axles are connected with the body by means of springs, dampers, and hydraulic devices in parallel [19]. Compared with passive suspension, active suspension adds hydraulic device to provide active control force. As shown in Fig. 4, when the body is moving, it is assumed that the tilting angle u and pitch angle h of the car body are very small. According to the motion law of rigid body, the displacements Zsfl; Zsfr; Zsrl; Zsrr of car body at each wheel are: Zsfl z bf h\u00fe 0:5au; Zsfr z bf h 0:5au; Zsrl z\u00fe brh\u00fe 0:5au; Zsrr z\u00fe brh 0:5au; \u00f01\u00de where z is the lift displacement of the center of gravity of car body; bf and br is the wheelbases from the center of gravity of the bodywork to the front and rear wheels, respectively; a is the tread of the left and right wheel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000792_s11668-015-0057-y-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000792_s11668-015-0057-y-Figure12-1.png",
        "caption": "Fig. 12 Finite element simulations of the failed gear: (a) model, (b) meshing, (c) stress distribution on the gear surface, and (d) stress distribution at cross-sectional area at the gear cross-sectional area",
        "texts": [
            " It also reveals that the wear fatigue pitting of the working tooth flank should not be attributed to roughness. For exhibiting the stress distribution on the tooth surface, FE method was used to analyze the stress distribution when the teeth engaged with each other. Accurate calculation of loads, deformations, and stresses in gear requires a numerical method such as the finite element method (FEM). In this paper, FE model was created and used to present the stress distribution in the gear. A solid model was created first, and then it was meshed with tetrahedral units (Fig. 12a, b). Properties of gear material are listed in Table 6. The simulation results are shown in Fig. 12c, d, including the stress distribution on the tooth flank and at the cross-sectional area. As shown in Fig. 12c, at least three teeth on the small gear are engaging with that of the big gear. The stress level of the engaging position is remarkably higher than the position without engagement. The maximum stress usually appears at the tooth flank near the pitch. Under normal conditions, the failure of the tooth flank near the pitch may occur earlier than other positions, because its stress level is higher than other positions. In this regard, the simulation results are closely in accordance with the features of the failed gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001889_tcyb.2015.2470225-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001889_tcyb.2015.2470225-Figure1-1.png",
        "caption": "Fig. 1. (a) Contour of the cost function f (x) and trajectory of the members of the group (black solid lines), where symbol \u201co\u201d indicates the initial position, and the arrows represent the direction of their trajectory. (b) Contour of the performance function and communication network in the group, where the arrows indicate the flow of the information, the thickness of the line indicates the strength in attraction between members, and the symbol \u201co\u201d marks the location of the individual at time step 150. (c) Cohesion \u03b3 (t) and performance J(t) of the group.",
        "texts": [
            " Recall that N denotes the number of members in the group. We choose xi \u2208 R 2 for all i = 1, . . . ,N, and f : R 2 \u2192 R to be a continuously differentiable function. We consider two different complex task environments: first, we choose a function that has three local minima, three local maxima, one global minimum x\u2217 such that f (x\u2217) = 0, and flat regions. The equation of this function is f (x1, x2) = 3(1 \u2212 x1) 2 exp [ \u2212 ( x2 1 ) \u2212 (x2 + 1)2 ] \u2212 10 (x1 5 \u2212 x3 1 \u2212 x5 2 ) exp ( \u2212x2 1 \u2212 x2 2 ) \u2212 1 3 exp [ \u2212(x1 + 1)2 \u2212 x2 2 ] + 6.5511. Fig. 1 shows a contour plot of this function. This function allows us to study the behavior of the group when the tasksolving process can lead to different solutions. In the second scenario, we choose a function defined by a plane. This function allows us to study the dynamics of the group in terms of performance and cohesion when the tasksolving process is continual, and does not have a fixed solution point. To quantify the group cohesion and performance at a given time t, we define three measurements: group cohesion in position, defined as \u03b3 (t) = \u2212 1 N(N \u2212 1)/2 \u2211 i,j\u2208V \u2225\u2225xi(t)\u2212 x j(t) \u2225\u22252 (39) is the negative sum of the pairwise distances between the individuals at time t",
            " To facilitate the analysis of the attraction patterns of the group, in this simulation there is no attraction in velocity, that is, b = 0. The communication network between the group members follows a symmetric ring topology (each node has two neighbors). This means that each group member communicates with his or her contiguous group members via a bidirectional flow of information. The discretization of the differential equations is done using Euler\u2019s approximation method with sampling time T = 0.1. Fig. 1 shows the behavior of the individuals at different time steps. From the trajectory of the agents in Fig. 1(a), we can observe that initially the individuals tend to go to their closest local minimum or saddle point. This is due to the fact that initially the force that drives each individual is more oriented toward solving the task than being attracted to the rest of the group. However, individuals start increasing their attraction to those that are connected to them and have a better relative performance as Fig. 1(b) illustrates. There is a point where the attraction weights become large enough that the force driving the individuals points toward those individuals with better relative performance, even if it implies moving in a direction that locally does not optimize its performance. Note that, since the agents are not fully interconnected, some individuals sequentially jump to minima until eventually they reach the global one. This behavior is evident in Fig. 1(c), where the group performance decreases at some points where group cohesion increases. A further interpretation of the mathematical model in (2)\u2013(8) allows us to explain the attraction patterns in the dynamics observed in Section IV. Using simple algebraic operations, we can rewrite the component associated with the attraction in position and velocity in (5) as ai(t) = \u2212w\u0304i(t) ( xi(t)\u2212 x\u0304i(t) )\u2212 bw\u0304i(t) ( vi(t)\u2212 v\u0304i(t) ) (42) where w\u0304i(t) = \u2211 j\u2208Ni wij(t) is the sum of the attraction in position strengths",
            " In a similar way, we can compute v\u0304i(t). During the process of solving the task, from (7), we have that an individual tends to be more attracted to the ones that are performing better than him or her, that is, have a higher contribution in the computation of the weight average w\u0304i(t). As some of the wij increase, term w\u0304i(t) in (42) becomes larger and the influence of the attraction component in force ui(t) becomes significantly larger than the component associated with performance optimization. In Fig. 1, we can observe that group members that initially move to local minima or saddle points later move toward those individuals that have better performance. For example, individual 2 moves to a saddle point, but later it gets attracted to 1, who is performing better. Also, individual 3 reaches a local minimum, but it later climbs up the valley because it gets attracted to individual 2 who is performing better (because 2 is being attracted to 1). These dynamics allow the group to eventually reach the global minimum and remain cohesive",
            " We simulate the model under different topologies of the network that define how the members of the group communicate between each other. First, we compute group cohesion and performance under three topologies: a wheel, in which most of the members of the group communicate with only one individual. A ring, in which every individual communicates with only two other group members. And a line, in which is similar to the ring topology but two of the group members communicate with only one individual. For each topology, the parameters of the model are chosen under three different scenarios: 1) as in the simulation in Fig. 1, where the performance optimization component of force ui is initially stronger than the attraction ones; 2) dynamics for the attraction weights and performance optimization strength (i.e., \u03b1w = 0 and \u03b1\u03b7 = 0), in which the attraction strength is significantly larger than that of performance optimization; and 3) again no dynamics for the attraction weights and performance optimization strength (i.e., \u03b1w = 0 and \u03b1\u03b7 = 0), but in this case the attraction and performance optimization strengths have a similar influence on the force ui"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001537_0954406218791636-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001537_0954406218791636-Figure14-1.png",
        "caption": "Figure 14. Representation of the turbine-generator shaft line layout at rated speed after alignment and correction of the distortion of pedestal #2 in the horizontal direction, by adjustment of bearing #2. Bearing #2 operating position has shifted back to yield optimum loading.",
        "texts": [
            " Such a distortion in multibearing shaft line would influence the operating eccentricity of the journal bearing as shown schematically in Figure 13 where the pedestal distortion results to the bearing #2 horizontal displacement of dz, 2. Then, the equilibrium position is shifted and the journal obtains (supposed to) an equilibrium closer to the vertical centreline of the bearing shell. The adjustable journal bearing is supposed to enable the corrective action to locate the bearing bore in the optimum position regardless the pedestal distortion, permitting the journal bearing operation as designed, see Figure 14. A distortion of the casing (stator) may result to further non-concentric rotor\u2013stator operation (additionally to the already eccentric operation of the rotor due to journal bearing clearance). The distorted casing centreline is shown in red in Figure 15, where the bore of bearings #2 and #3 are not anymore concentric to the casing. The corrective action of the adjustable bearings can retain the concentricity of the bearing bore and the casing if the casing distortion is within some extent, see Figure 16"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002267_tmag.2013.2281460-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002267_tmag.2013.2281460-Figure8-1.png",
        "caption": "Fig. 8. Iron loss distribution. (a) E&S model in [8]. (b) Proposed complexvalued model.",
        "texts": [
            "5 kW three-phase induction motor model is taken as another example, in which the pole number is two, the stator slot number is 30 and rotor is 26, as shown in Fig. 6(a). The motor runs at no-load and to make the mesh of stator fixed when rotor rotates an edge line at air region is defined, as shown in Fig. 6(b). The B loci distributions in stator core calculated from the proposed complex-valued model are shown in Fig. 7. From this figure, it is can be observed that the NGO electrical steel is magnetized with a lot of rotational magnetic flux densities, which are different from the transformer core. Fig. 8 shows the iron loss distribution in stator core. In Fig. 8, the distribution rules calculated from proposed complex-valued and E&S models in [8] are almost same, in which the maximum loss appears at the bottom of the slot near the middle part of stator core where the magnetic force line concentrates and goes along the transverse direction. The total loss can be calculated as follows: Ptotal = \u03c1 \u00b7 Dp \u00b7 Nos \u00b7 Nes\u2211 i=1 Pi loss \u00b7 Si (W) (10) where Ploss is from (3), Si is the area of each mesh element, Dp is the volume of each sheet, and Nos is the number of laminated sheets"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure10-1.png",
        "caption": "Figure 10: a) von-Mises stress and b) stress in xx-direction in sectional view 2 for the flux of force adapted lattice structure",
        "texts": [
            " Figure 9 shows the simulation result for the z-displacement of the structure. In order to make the results more comparable, the colour scale is the same as in figure 5. The displacement in the point of force application is 0.139 mm, which means a slight improvement compared to the periodic structure. Together with the applied force of 300 N, a stiffness of 2166 N/mm results for the structure. Due to the extensive reduction of the structure\u2019s mass, the ratio of stiffness to mass improves for 81 % towards a value of 25.9 N/(mm*g). Maximum Stress. In figure 10, the von-Mises comparison stress and the stress in xx-direction in cutting plane 2 are depicted for the flux of force adapted lattice structure. Figure 10 b) shows the stress in xx-direction for the cross section of the struts in cutting plane 2 (see figure 3). In order to make the results more comparable, the same colour scale has been used as in figure 6 b). It can be recognized, that only minor stress variations appear inside the single struts. This leads to the conclusion, that the bending loads in this sectional view have severely been reduced. Nevertheless, the maximum value for the von-Mises stress has risen to 74.63 N/mm\u00b2 (see figure 10 a). Under the assumption, that this maximum stress linearly depends on the applied force, a maximum load of 221.1 N can be calculated for the obtaining of the limit of elasticity. This means a worsening of 46 % compared to the periodic structure. The reason for the increased maximum stress can be seen in figure 11. The figure shows the sectional view at cutting plane 1 (see figure 3). It can be recognized, that there appears severe uparching, especially in longer struts with strong curvature. This uparching leads to high stresses along the struts and especially to stress peaks near the nodes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure4-1.png",
        "caption": "Figure 4: Geometry, constraints and load for periodic lattice structure",
        "texts": [
            " This is the currently established approach for the application of lattice structures in lightweight design. Several software tools for the generation of such geometries for additive layer manufacturing are already available. Structure Geometry. The structure is created by joining one and the same cubic elementary cell in the three directions in space. As mentioned before, a strut diameter of 2 mm and an interval of 5 mm have been chosen. The explained build-up has been applied on the design space in figure 3 (see figure 4). This leads to a total mass of 147.5 g for the whole structure. The bearing and the force application are realized as described before. Stiffness. Figure 5 shows the simulation result for the z-displacement of the structure in figure 4. The displacement in the point of force application, which is the reference point to compare the different structures, is 0.142 mm. Together with the applied force of 300 N, a stiffness of 2117 N/mm results for the structure. To make the performance more comparable to other structures, the ratio of stiffness to mass is calculated. Here, a value of 14.4 N/(mm*g) arises. Maximum Stress. In figure 6 a), the von-Mises comparison stress is depicted for the regular lattice structure under the given load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000456_j.neucom.2012.11.007-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000456_j.neucom.2012.11.007-Figure1-1.png",
        "caption": "Fig. 1. Spherical inverted pendulum.",
        "texts": [
            " In this paper, a scheme based on neural network approximation of the feedforward function without solving the regulator equations approximately will be adopted. Since the dimension of the feedforward function is only equal to two, this new scheme is much simpler than the existing approaches. Moreover, when all the states are available, our design offers certain robustness to plant parameter variations and leads to good tracking performance. & 2012 Elsevier B.V. All rights reserved. 1. Introduction The spherical inverted pendulum can be shown in Fig. 1 [12,15] where x,yAR represent the position of the base of the pendulum in the horizontal plane and X,YAR represent the x and y positions of the vertical projection of the center of the pendulum onto the horizontal plane, Fx,FyAR are the control forces being applied to the cart at the base of the pendulum, m is the mass of the uniform rod, L is the distance from the base of the pendulum to the center of mass, and g is the gravitational constant. The motion equations of the spherical inverted pendulum are as follows [12,15]: _x \u00bc f \u00f0x\u00de\u00feg\u00f0x\u00deu y\u00bc h\u00f0x\u00de \u00f01\u00de where x1 \u00bc x, x2 \u00bc _x, x3 \u00bc y, x4 \u00bc _y, z1 \u00bc X, z2 \u00bc _X , z3 \u00bc Y , z4 \u00bc _Y u1 \u00bc Fx, u2 \u00bc Fy, u\u00bc \u00bdu1 u2 T x \u00bc \u00bdx1 x2 x3 x4 z1 z2 z3 z4 T ll rights reserved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000685_58578-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000685_58578-Figure3-1.png",
        "caption": "Figure 3. Two configurations of the horizontal thrusters",
        "texts": [
            "Thus (14) can be written as: ( ) ( ) ( ) ( ) ( ) ( ) ( ) 1 4 4 4 4 mE n E n E n E n E n E n \u03c9 \u03c3 \u03b8\u03b5 \u03bc \u03bc \u03bc \u03bc+ \u2248 + + + < + + + = (14) According to (8) and (14), the tracking errors for the proposed petri-net-based dynamic RFNN controller will converge to zero gradually. 3. Fault Tolerable Thrust Allocation 3.1 Model of Thrust Control Allocation ROV can enable diving with only one thruster in the vertical direction, if enough load is available. Therefore, we focus our research on fault tolerable thrust allocation in the horizontal plane. Usually, an open frame underwater vehicle has four horizontal thrusters, denoted iHT, i=1,2\u20264. There are two common configurations of the horizontal thrusters, e.g., X-shaped and Cross-shaped configurations (Figure 3) [10]. Their position and orientation are illustrated in Table 2 and Table 3. Therefore, the vectors of forces and moments, exerted by HT, can be expressed as: ( ) ( ) ( ) ii iHT i ii i i T i ix iy i i i ix y FeF F r eM e e r e r e F \u03c4    = =    \u00d7     = \u00d7 \u00d7            5Huang Hai, Wan Lei, Chang Wen-tian, Pang Yong-jie and Jiang Shu-qiang: A Fault-tolerable Control Scheme for an Open-frame Underwater Vehicle where iF  is the force exerted by thruster i on the ROV, ( )i i r i iM F r e= \u00d7    is the moment deduced from iF  ",
            " Operating the heading control of SY-II is much harder because of the length-width ratio of the open-frame of the underwater vehicle and its strong non-linear and random disturbance from the tether. In the heading control experiments of Figure 7, we compared the systems in four situations: (1) PRFNN controller of section 2 without thrusters fault; (2) PRFNN controller with bi-criteria FTC ( 0.5,a = 510\u03b3 = , where a has been selected from heading control simulations and experiments, for the two criteria of convergence speed and control errors, \u03b3 should be set as large as possible to allow for computer simulations, the left main thruster (1HT in Figure 3.(b)) 70% fault and tail side thruster (4HT in Figure 3.(b)) 100% fault); (3) PRFNN controller with 2- norm FTC ( 1,a = 510\u03b3 = ; the left main thruster (1HT in Figure 3.(b)) 70% fault and tail side thruster (4HT in Figure 3.(b)) 100% fault), which is similar to a pseudo inverse solution in [11] and [12]; (4) FNN controller mass(kg) xI 2( )Nms yI 2( )Nms zI 2( )Nms xyI 2( )Nms xzI 2( )Nms xzI 2( )Nms 111.9 97.3 26.1 56.8 0 0 0 Table 4. SY-II inertial parameters Dimensionless coefficient x y z r q p First order terms 27.65 47.26 44.86 62.63 57.55 141.33 Table 5. SY-II Hydrodynamic parameters 9Huang Hai, Wan Lei, Chang Wen-tian, Pang Yong-jie and Jiang Shu-qiang: A Fault-tolerable Control Scheme for an Open-frame Underwater Vehicle without any faulty thrusters",
            " When an infinity-norm optimal solution is introduced through weighting factor a, the FTC is transformed into a bicriteria FTC scheme. The bi-criteria FTC scheme behaves with better convergence results in comparison with the 2-norm FTC scheme during the heading control experiments, because infinity-norm optimization enables a better thrust magnitude optimization and a self continuity solution. Cruising simulations and experiments have been carried out with the precondition that the left main thruster (1HT in Figure 3.(b)) has 40% fault and the tail side thruster (4HT in Figure 3.(b)) has 100% fault. In the cruising simulation under or without a current effect (eastward 0.2m/s) in Figure 8 (a) and (b), the desired motion trajectory is (0,0)-(70,200)-(0,300)-(120,300)-(120,200)-(0,0). Furthermore, cruising experiments have been conducted for a SY-II equipped with an Ultrosonic Doppler Velocity Meter in Figure 6 (a) and 8 (c), the desired motion trajectory is (-7,-20)-(-2,-2)-(5,-2)-(5,-22)-(-7,-20). From Figure (8), the bi-criteria FTC outperforms the 2-norm FTC for fault tolerance control in reducing overshoot during yawing, approaching the desired trajectory and cruising against the current"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003949_s00773-019-00660-1-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003949_s00773-019-00660-1-Figure5-1.png",
        "caption": "Fig. 5 Definition of encounter angle [18]",
        "texts": [
            ",\ud835\udf14y, vx, vz, rpx, rpz,mb) T for underwater glider in vertical plan: where ucx, vcz are the components of ocean currents in x\u2013z plane, while wave is the wave force defined in Eq.\u00a042. The rest of the terms used in Eqs.\u00a028\u201330 are as follows: The direction of the wave force depends upon the direction of wave velocity with respect to encounter angle which is based on the position control of AUVs and ships [18]. The direction and orientation of the encounter angle with respect to the glider is shown in Fig.\u00a05. The dynamic equations of wave and current forces are added to the dynamic model of the AUG. The wave forces depend on the direction of the wind and encounter angle, . In case of following sea = 0\u25e6 , the wave (25)x\u0307 = vx cos \ud835\udf03 + vz sin \ud835\udf03, (26)z\u0307 = \u2212vx sin \ud835\udf03 + vz cos \ud835\udf03, (27)?\u0307? = \ud835\udf14y, (28)?\u0307?y = 1 Jy\ufffdF\ufffd \u239b \u239c\u239c\u239d rpxaz m3 X3 \u2212 rpzaz m1 X1 \u2212rpxax(\ud835\udf14yr\u0307px \u2212 w3) \u2212rpzaz(\ud835\udf14yr\u0307pz \u2212 w1) + axazY \u239e \u239f\u239f\u23a0 , (29)v\u0307x = 1 m1\ufffdF\ufffd \u239b \u239c\u239c\u239c\u239d \u2212 rpzaz Jy Y + dz m\u0304 X1 \u2212 c m3 X3 + c \ufffd \ud835\udf14yr\u0307px \u2212 w3 \ufffd \u2212 \ufffd az + bx \ufffd\ufffd \ud835\udf14yr\u0307pz + w1 \ufffd \u00b1 ucx \u00b1 \ud835\udf02wave \u239e \u239f\u239f\u239f\u23a0 , (30)v\u0307z = 1 m3\ufffdF\ufffd \u239b\u239c\u239c\u239c\u239d rpxax Jy Y \u2212 c m1 X1 + dx m\u0304 X3 +(ax + bz)(\ud835\udf14yr\u0307px \u2212 w3) \u2212c(\ud835\udf14yr\u0307pz + w1) \u00b1 ucz \u00b1 \ud835\udf02wave \u239e\u239f\u239f\u239f\u23a0 , ax = 1 m\u0304 + 1 m1 , az = 1 m\u0304 + 1 m3 , bx = r2 px Jy , bz = r2 pz Jy , c = rpxrpz Jy , dx = ax + bxax + bz, dz = az + bzaz + bx, X1 = \u2212m3vz\ud835\udf14y \u2212 m\u0304(vz + r\u0307pz \u2212 rpx\ud835\udf14y)\ud835\udf14y \u2212 m0g sin \ud835\udf03 + L sin \ud835\udefc \u2212 D cos \ud835\udefc, X3 = m1vx\ud835\udf14y + m\u0304(vx + r\u0307px + rpz\ud835\udf14y)\ud835\udf14y + m0g cos \ud835\udf03 \u2212 L cos \ud835\udefc \u2212 D sin \ud835\udefc, Y = (m3 \u2212 m1)vxvz \u2212 m\u0304g(rpx cos \ud835\udf03 + rpz sin \ud835\udf03) +MDL. forces are added to the dynamic model, while in case of head sea = 180\u25e6 , the wave forces are subtracted from the model, as shown in the Fig.\u00a05. The equations of motions (Eqs.\u00a025\u201330) describe the motion of the glider in vertical plane [10]. The desired output variables are the values with the subscript \u201cd\u201d. The objective of the control action is to reduce the differences for the internal actuator\u2019s inputs rpx, rpz and mb. The desired path for shifting mass rpxd(t), rpzd(t), and desired path for variable ballast mass control mbd(t) are time varying, which are defined in Eqs.\u00a032 and 33: where Vd = \u221a( v2 x + v2 z ) where mf1 and mf3 are the added masses defined by an added mass matrix, Mf = diag(mf1,mf2,mf3) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000209_dscc2013-3740-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000209_dscc2013-3740-Figure2-1.png",
        "caption": "Figure 2. THE STRUCTURE OF THE CONICAL SHAPE MANIPULATOR DRIVEN BY TWO CABLES. PULLING THE CABLE T1, THE POINT LOAD \u2212T1tc1 AND THE DISTRIBUTED LOAD w1(s) GENERATE THE BENDING MOVEMENT.",
        "texts": [
            "org/about-asme/terms-of-use can find the tip position of the manipulator with respect to the base frame (x, y, z): xee = \u222b L 0 (1 + q(s))tx(s) ds (5) yee = \u222b L 0 (1 + q(s))ty(s) ds (6) zee = \u222b L 0 (1 + q(s))tz(s) ds (7) where L is the arm length. In the following section we describe the statics model of the conical shape manipulator, i.e. the relation between the cable tensions T1 e T2 and the deformation variables (k(s), q(s)). Kinetics of the conical shape manipulator The conical shape manipulator has two cables embedded in the structure (Fig. 2). The cables, which are coplanar with the midline, exert a stress on the structure realizing the deformations k(s) and q(s) (\u03be(s) = 0, \u03c4(s) = 0) in 2-D space. Unlike the existing work [8], we consider here the cables parallel to the midline, that is the cables maintain the same distance yc from the midline along the structure (Fig. 2). This geometric change simplifies the mechanical model equations, without modifying the manipulator property of realizing a spiral-like configuration. The positions of the two cables are defined as follows: uc1(s) = u(s) + yc(s)n(s) (8) uc2(s) = u(s)\u2212 yc(s)n(s) (9) Moreover, we introduce the radius R(s) of the section s: R(s) = ( Rmin \u2212Rmax L ) s+Rmax (10) When the cables are pulled, they exert a point load at the spot where they are anchored, and a distributed load along the cable length. The point load is equal in magnitude to the cable tension and tangent to it \u2212Titci(s)(Ti > 0) (Fig.2). The distributed load is centripetal and proportional to the curvature of the cable wi(s) = Tidtci/dSc [9] (Fig.2). Where i \u2208 {1, 2} identifies one of the two cables and Sc represents the arclength parametrization of the cable. Since the cables are immersed inside the body, the statics problem is a following force problem. The equilibrium equations of our manipulator are: mi(s) = \u2212Titci(L)\u00d7 [u(s)\u2212 uci(L)] + + \u222b L s Ti dtci d\u03c3 \u00d7 [u(s)\u2212 uci(\u03c3)] d\u03c3 (11) li(s) = \u2212Titci(L) \u00b7 t(s) + +t(s) \u00b7 \u222b L s Ti dtci d\u03c3 d\u03c3 (12) where li(s) and mi(s) are respectively the internal longitudinal stress and the internal torque of the robot arm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002052_isgteurope.2014.7028952-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002052_isgteurope.2014.7028952-Figure2-1.png",
        "caption": "Figure 2. Cross-section of G(Z)TACSR conductor.",
        "texts": [
            " - ZTACIR (Heat Resistant Aluminum Alloy Conductor Invar Reinforced) 22.75 mm in diameter [20-22]. - ACSS (Aluminum Conductor Steel Supported) 20.9 mm in diameter [23-24]. HTLS conductors are electrically and geometrically very similar to the conventional ACSR conductor. The main difference lays on the lower heat expansion coefficient. The direct consequence is that the conductors accept a higher temperature, for the same sag, with an increase of the ampacity of the line [25]. Gap-type conductors are described in [26]. \u201cFig. 2\u201d shows the cross-section of a G(Z)TACSR conductor. There is a small gap between the steel core and the innermost aluminium layer, in order to allow the conductor to be tensioned on the steel core only. This effectively fixes the conductor\u2019s knee-point to the erection temperature, allowing the low-sag properties of the steel core to be exploited over a greater temperature range. The gap is filled with heat-resistant grease to reduce friction between the steel core and the aluminium layer, and to prevent water penetration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003765_iros.2018.8594316-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003765_iros.2018.8594316-Figure9-1.png",
        "caption": "Fig. 9: Sensor arrangement.",
        "texts": [
            " Although the pair of index finger and middle finger, that of ring finger and little finger and the variable rigidity mechanism mentioned in the previous section are respectively moved by one actuator because of the limitation of the number of actuators, by knots in the branch point of palmar interossei muscle and dorsal interossei muscle and ring-shaped tendon connectors[11] in the other branch points. It should be noted that all muscles in this hand are placed in forearm although interossei muscles and the adductor pollicis muscle of human beings are intrinsic muscles because of the limitation of the size of actuators and the hand. 2) Sensor Arrangement: Nine loadcells are inserted in this hand. The position of loadcells is as shown in Fig. 9. Four loadcells are placed in the position of red circles in the palm. The others are placed in fingertips. The cables of loadcells in fingertips pass in coils of springs and all cables are connected to an amplifier board in the back of the hand. This hand can detect the touch to the object and measure the distribution of external force in the palm by referencing sensor values. 3) Exterior: This hand has exterior made of rubber to secure the softness and frictional force. An urethane sponge coated by urethane rubber (A60) is pasted on the palm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003135_0954406217720823-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003135_0954406217720823-Figure2-1.png",
        "caption": "Figure 2. Schematic of the initial stress.",
        "texts": [
            " Based on the Taylor series expansion, the shear stress can be written as follows \"r \u00bc 1 R @ @ \u00fe 1 R2 @w @ zop \" z \u00bc 1 R @ @ \u00fe 1 R2 @w @ zin \u00f08\u00de It should be noted that the out-of-plane deformations were fully considered (including the shear strain caused by the out-of-plane torsional and bending vibration) when the shear strain was derived. However, the shear strain caused by the in-plane deformations was ignored. The Hamilton principle is used to derive the equations of motion. The Hamilton principle can be expressed as follows Z t2 t1 U T W\u00bd dt \u00bc 0 \u00f09\u00de where U is the potential energy, T is the kinetic energy, and Wis the external force energy (including the internal pressure). Loads. To obtain the potential energy and external force energy, the loads acting on the ring should be given. As observed in Figure 2, considering the initial stress caused by the uniform pressure acting on the inner wall of ring, 0 denotes the initial stress given by the following 0 A \u00bc 1 2 Z 0 A 2r\u00fe p0b sin rd \u00bc p0br\u00fe Ar 2 2 \u00f010\u00de where 0 is the initial stress, A is the ring section, is the density, is the rotational speed, b is the belt effective width, and p0 is the internal pressure. The radial, tangential constraining forces, and moments are also included to model real engineering structures, as shown in Figure 3. Energies in the Hamilton principle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000025_1.3667502-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000025_1.3667502-Figure4-1.png",
        "caption": "Fig. 4 Geometry for filleted notch problem",
        "texts": [
            " With these changes equation (20) becomes <re = aETo [ ( f l + f)-2 fQ r {R + p)f(p)dp - / ( f ) ] (21) Notch Problem Introduction. T h e theoretical da ta in the preceding sections provide information on the general characters of the stress fields on the straight edges and a t the corner of a filleted rectangular notch. Those results, together with a s tudy of the temperature field generated during thermal shock, are combined in this section to provide an analysis of the stress induced a t a filleted notch by a thermal shock field specified, as shown in Fig. 4. The boundary temperature To is less than, or equal to, the largest possible difference between the temperature in the shocking fluid and the initial temperature of the elastic 3 / , space occupied by the notched plate. In Fig. 4, f is the distance meas- Journal of Engineering for Industry A U G U S T 1 9 6 2 / 3 4 5 Downloaded From: http://manufacturingscience.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmsefk/27464/ on 06/12/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use HALF PLANE EDGE o I / HOLE BOUNDARY CIRCLE T T(y)= Tag(y) T(r)= Tof(r) a) HALF PLANE AND HOLE ured into the plate f rom the boundary, either normal to t he straight edge or radially in t he fillet region. Analysis. A comparison of the stresses parallel to the straight and curved boundaries may be obtained from the rat io of equations (7) and (21), o'e/o'x = f i f ) - (R + f)~ ",
            " I t will now be shown t h a t when a proper limitation is placed upon the penetration distance s of the temperature field little chauge occurs in the character of the stress field in the region of the corner fillet compared to t h a t in a half plane. Tor an axisymmetric temperature field b2T/df2 + (R + f)~\\DT/b?) = (1 /D)(dT/dd) (23) while for a uniaxial Cartesian field, b2T/df2 = (l/D)i>T/bd (24) For the purpose of this demonstration assume t h a t a t a given 9, T = T0e-cr (25) where c > 1 to insure rapid decay of T in the f-direction. Some justification for the choice m a y be found in an article by Heisler [5]. Then referring to Fig. 4 {R + ?)-'(c)27c>F) = (R + f)~\\bT/by) = -cT0e-\u00b0\\R + f)~l (26) d2T/bf2 = (<)2T/by2) = c2Toe\"\" (27) and therefore the ratio of the two terms, defined as (3, is (R + f)-'(ar/Sr) d'T/dr2 c(R + f) (28) In order to set a mathematical limit to the penetrat ion distance, assume t h a t cr = cs = 4 so t h a t T/To = c~4 a t s = mR where m is a number less than 1. Then c = 4 / m R and jS = mR 4 (R + mR) 4(1 + m) (29) Then it follows t h a t for m in the order of l / t to l/i (or s ^ (R/2)) the term (R + f ) ~l(dT/dr) would be significantly less than Z>2T/dr2 and the temperature fields for the infinite half plane and the fillet would not differ significantly under thermal shock"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000288_amr.576.471-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000288_amr.576.471-Figure2-1.png",
        "caption": "Fig. 2: Effect of kenaf fiber loading in hybrid composite on impact strength",
        "texts": [
            " However, 0/30 KG composite showed higher flexural strength compared to 30/0 KG composite material. When the specimen is loaded in flexural state, both outer sides are given high stress due to large compression and tension. As matter of fact, fibre orientation was found to have a significant role in the properties of kenaf fibre composite. The 22.5/7.5 KG showed lower strength followed by 30/0 KG composite. Therefore, it can be said that kenaf from a green source is a potential fibre for hybridization towards the green composite with 15/15 KG in the reinforcement. Impact Properties. Fig. 2 illustrates the measured impact strength values of the hybrid composites in this investigation. It is seen from Figure 2 that the impact strength of the composite increase gradually with the increase of volume fraction of untreated kenaf in the reinforcement, however, treated kenaf hybrid composite showed lower impact strength compared to untreated one. This declination of the impact energy could be due to the lack of proper lamination between reinforcement and matrix materials during preparation and mixing of resin and hardener. Highest impact strength was observed for the untreated 15/15 KG (kenaf and glass) hybrid composite compared to treated hybrid The effects of kenaf on the flexural and impact properties of reinforced unsaturated polyester composite (UPE) have been studies and following conclusions can be drawn: 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003816_s12541-019-00014-2-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003816_s12541-019-00014-2-Figure9-1.png",
        "caption": "Fig. 9 Parts of the blade bearing (left) and connection graph of the two sub-assemblies (right)",
        "texts": [
            " Because the leak between the blade foot and hub can occur at any place of the contact surface, in this study, the maximum clearance dmax between the contact (12)A12 = A2 \u2212 A1 (13) { gmax = max ( A12 ) gmin = min ( A12 ) (14)g(T) = gmax + dg\u2212w(T) = max ( A12 ) + dg\u2212w(T) surfaces is set to represent the sealing performance. However, minimum clearance dmin between the contact surfaces represents the flexibility of controlling the pitch, i.e., a larger dmin corresponds to a better flexible performance. Figure\u00a09 (left) and Table\u00a03 show the related tolerances of the parts of assembly, and the gap between the blade foot and hub is determined by the dimension tolerances X3 of hub, X2 of the blade carrier and X1 of blade foot as well as their corresponding geometric tolerances t1 , t2 , t3 and t4 . Figure\u00a09 (right) displays the connection graph of this assembly, exhibiting that there are two functional requirements (FRs), which are forming the gap. The fit of the blade foot and blade carrier is perfect, and here CFE3 is null. The related tolerances parameters in Fig.\u00a09 (left) are listed in Table\u00a04. The detailed calculation process of the assembly deviation analysis of this assembly was presented in a previous paper [32]. The calculation results of the assembly deviation with only considering the dimensional and geometric tolerances are presented in Table\u00a05. In this section, the effect of the deformation on the assembly clearance is discussed. The related range of the temperature is (25\u201360\u00a0\u00b0C) and rotation speed is (60\u2013210 r/min), and the calculation results of second part dg\u2212r of the assembly gap are presented in Tables\u00a05 and 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003467_s40997-018-0184-7-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003467_s40997-018-0184-7-Figure3-1.png",
        "caption": "Fig. 3 The coordinate system describing the gear machining process with rack cutter",
        "texts": [
            " The parabolic equation in coordinate system obxbyb is described as rb \u00bc u au2 0 1 2 664 3 775 \u00f01\u00de By coordinate transformation, the normal tooth surface equation of modified rack cutter in coordinate system oc1xc1yc1zc1 is described as rc1 \u00bc xc1 yc1 zc1 1 2 664 3 775 \u00bc Mc1brb \u00f02\u00de where Mc1b \u00bc 1 0 0 0 0 1 0 am 0 0 1 0 0 0 0 1 2 6664 3 7775 cos a sin a 0 0 sin a cos a 0 0 0 0 1 0 0 0 0 1 2 6664 3 7775 1 0 0 dp 0 1 0 0 0 0 1 0 0 0 0 1 2 6664 3 7775 By coordinate transformation, the transverse tooth surface equation of modified rack cutter in coordinate system ocxcyczc is described as rc \u00bc xc yc zc 1 2 664 3 775 \u00bc Mcc1rb \u00f03\u00de where Mcc1 \u00bc 1 0 0 0 0 cos b sinb 0 0 sin b cos b 0 0 0 0 1 2 664 3 775 1 0 0 0 0 1 0 0 0 0 1 l 0 0 0 1 2 664 3 775 The coordinate system describing the gear machining process with rack cutter is shown in Fig. 3. By coordinate transformation, the tooth surface equation of pinion/gear in coordinate system oixiyizi is described as ri \u00bc xi yi zi 1 2 664 3 775 \u00bc Mcoirc \u00f04\u00de where Mcoi \u00bc cos/i sin/i 0 0 sin/i cos/i 0 0 0 0 1 0 0 0 0 1 2 664 3 775 1 0 0 ri 0 1 0 si 0 0 1 0 0 0 0 1 2 664 3 775; /i \u00bc si=ri According to the Willis law, the rack cutter tooth surface envelops the gear tooth surface in the process of machining; this process must satisfy the following meshing equation: f \u00f0ui; li;/i\u00de \u00bc n~ci v~\u00f0ci;i\u00de ci \u00bc 0 \u00f05\u00de By solving formula (5), li can be replaced by ui and /i, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003766_rcs.1983-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003766_rcs.1983-Figure4-1.png",
        "caption": "FIGURE 4 Schematic model of the grasper",
        "texts": [
            " JT q is the Jacobian matrix transform of generalized velocity q\u0307 into angular velocity ?\u0307?. It is given by time differentiate Equation 3 and is expressed by following equation. ?\u0307? = Jqq\u0307 (15) Jq = [ \u2212\ud835\udf03 sin \ud835\udeff cos \ud835\udeff \ud835\udf03 cos \ud835\udeff sin \ud835\udeff ] (16) As shown in Section 2.2, the grasping force is estimated using a grasper mechanism model and an equation of motion. The grasper has a groove cam mechanism and is driven by the cylinder via a push-pull wire. A schematic model of the grasper and its geometric parameters are shown in Figure 4 and Table 1. a and b are perpendicular to the axis of the upper and the lower groove. c shows the length from the intersection point to the maximum cylinder position Xgmax . \ud835\udf03cam represents the inclination angle of the upper groove. d shows the length from the cam to the center of rotation of the grasper, which depends on the cylinder position Xg . The relationship between the grasper angle \ud835\udf03g and the cylinder position Xg can be obtained geometrically as follows: \ud835\udf03g = arcsin b d + arcsin a d \u2212 \ud835\udf03cam\u2216textbackslash (17) d = \u221a b2 + (c + Xgmax \u2212 Xg)2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001801_978-90-481-9707-1_108-Figure50.4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001801_978-90-481-9707-1_108-Figure50.4-1.png",
        "caption": "Fig. 50.4 Mechanical model of quadrotor with single payload",
        "texts": [
            " The formulation of the tangential acceleration a and the normal acceleration an can be obtained, respectively, using the theorem of motion of center of mass law of conservation of momentum, which are calculated through the suspension angles \u02dbx and \u02dby . Then, the forces caused by the payload in respective axes are derived as follows: Fox D MLaox D MLa x cos\u02dbx MLanx sin \u02dbx D ML R\u0328xL cos\u02dby cos\u02dbx ML P\u0328 2xL cos\u02dby sin \u02dbx Foy D MLaoy D MLa y cos\u02dby MLany sin \u02dby D ML R\u0328yL cos\u02dbx cos\u02dby ML P\u0328 2yL cos\u02dbx sin \u02dby Foz D MLaoz MLg D ML R\u0328xL cos\u02dby sin \u02dbx CML P\u03282xL cos\u02dby cos\u02dbx CML R\u0328yL cos\u02dbx sin\u02dby CML P\u0328 2yL cos\u02dbx cos\u02dby MLg Using the Newton-Euler function, the mechanical model reflecting the force acting on the quadrotor is shown in Fig. 50.4. Therefore, the model of the quadrotor with payload can be expressed as .M CML/ Rx D U1.cos sin cos C sin sin / Fox .M CML/ Ry D U1.sin sin cos cos sin / Foy .M CML/Rz D U1.cos cos / Mg C Foz R D P P Jx Jz Jx C l Jx U2 R D P P Jz Jx Jy C l Jy U3 R D P P Jx Jy Jz C 1 Jz U4 (50.13) In order to gain insight into the dynamics of the quadrotor with slung payload, the numerical simulation for the slung payload with quadrotor is presented in this section. The simulation parameters for the quadrotor are chosen as M D 1:52 kg, and its moments of inertia were estimated as 0:03, 0:03, and 0:04 kg m2 for the x, y, and z body axes, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000851_s12541-015-0332-6-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000851_s12541-015-0332-6-Figure3-1.png",
        "caption": "Fig. 3 CFRP grinding wheels employed for crankshaft, camshaft, and gear shaft grinding51 (Printed with permission from Asen)",
        "texts": [
            " Based on the mass of research on the advantages of grinding with CFRP wheels, Mach Rotec developed a CFRP grinding wheel for supersonic speed grinding (i.e., grinding speeds greater than 333 m/s).51 Compared to steel substrate grinding wheels, this grinding wheel was 90% lighter and had excellent stability and damping characteristics. CFRP grinding wheels manufactured by this company have been successfully applied to the grinding of precision parts like camshafts, gear shafts, crankshafts, etc. (Fig. 3). Compared to steel substrate grinding wheels, this CFRP grinding wheel has a higher material removal rate, along with a 25% reduction in machining time for the same conditions, thereby resulting in improved production capacity. Particularly for machining precision parts like cams, crankshafts, gear shafts, etc., grinding speeds up to 180~230 m/s can be used, with a normal grinding force between 40 and 300 N. Owing to the significant static and dynamic characteristics, the quality of the machining surface was measured; the results showed that Ra < 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001928_icit.2015.7125147-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001928_icit.2015.7125147-Figure9-1.png",
        "caption": "Fig. 9. Robot diagram.",
        "texts": [
            " For example, (18) shows the force Fx(r) in x direction. U(r) = U(rx) + U(ry) + U(rz) (17) Fx(r) = 12\u03f5 \u03c3 [( rx \u2212 rx0 \u03c3 + 1 )\u221213 \u2212 ( rx \u2212 rx0 \u03c3 + 1 )\u22127 ] (18) The image of 3D L-J potential field and it\u2019s stream plot are shown in Fig. 8. The experiment was conducted to confirm the validity of stereo image reconstruction and proposed modulated potential field will fulfill the requested performance for position adjusting with human interaction. A three dimensional parallel link manipulator was used in this experiment. Fig. 9 shows diagram of the robot, and Fig. 10 shows the experimental setup. It is consisted of two DC motors (Maxon R20) and a direct drive rotary motor (SGMCS-07B). A stereo camera (scanv2pro) was mounted on the joint1. The Jacobian matrix of this 3D parallel link manipulator is shown in (19). J = \u2212L1 sin \u03b81 \u2212L2 sin \u03b82 0 L1 cos \u03b81 cos \u03b83 L2 cos \u03b82 cos \u03b83 \u2212(L1 sin \u03b81+L2 sin \u03b82) sin \u03b83 L1 cos \u03b81 sin \u03b83 L2 cos \u03b82 sin \u03b83 (L1 sin \u03b81+L2 sin \u03b82) cos \u03b83  (19) The relation of motor torque \u03c4 and the force of end effector F are shown in (20)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001800_ijaac.2018.095109-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001800_ijaac.2018.095109-Figure1-1.png",
        "caption": "Figure 1 The schematic diagram of the TRMS",
        "texts": [
            " But here angle of attack is fixed due to its two-dimensional nature. Therefore, the DC motor input is changed to modify rotors speed. Two tachometers, mounted on these two rotors are used to measure the velocity of the DC motors and one position sensor is pivoted on the beam to sense rotors position. Electrical unit transfers the measured signal from the tachometers and the sensor to a computer and allows the control signal to the TRMS through an input-output port. The schematic diagram of TRMS is shown in Figure 1 (TRMS 33-220 user manual, 1998; TRMS advanced teaching manual, 1998). In this present work the TRMS model is controlled in three different modes such as: a 1-DOF pitch rotor control, where cross coupling dynamics is ignored and only the main rotor is controlled b 1-DOF yaw rotor control, where cross coupling dynamics is again ignored and only the tail rotor is controlled c 2-DOF rotor control, where both the rotors are controlled simultaneously and all coupling effects are included here. In all modes of operations the control objective is set the beam of the TRMS according to the desired or reference trajectory by controlling the pitch angle and yaw angle with minimum possible transient error"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002416_j.neucom.2015.11.054-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002416_j.neucom.2015.11.054-Figure1-1.png",
        "caption": "Fig. 1. Model of each agent in a MASS.",
        "texts": [
            " Example 1 In [32] the model for the Multi-Agent Supporting Systems (MASS) with each agent supporting by one pillar was introduced and it was shown that the MASS has potential applications in earthquake damage-preventing buildings, water-floating plants and large-diameter parabolic antennae or telescopes. It was shown that each agent in a MASS can only be modeled by a singular system when this MASS consists of many independent blocks and each block is supported by several pillars. for the case that each agent in a MASS is supported by two pillars called Unit I and Unit II respectively, as shown in Fig. 1, where m the mass, d the damping coefficient and k the stiffness coefficient, let xiI t\u00f0 \u00de; xiII t\u00f0 \u00de;\u03c5iI t\u00f0 \u00de;\u03c5iII t\u00f0 \u00de denote the heights and velocities of Unit I and Unit II respectively, then agent i\u00f0iA\u00f01;\u22ef;N\u00de\u00de can be described by E _xi t\u00f0 \u00de \u00bc Axi t\u00f0 \u00de\u00feBui t\u00f0 \u00de \u00f038\u00de gn for nonlinear singular systems, Neurocomputing (2015), http: where xi t\u00f0 \u00de \u00bc xiI t\u00f0 \u00de \u03c5iI t\u00f0 \u00de xiII t\u00f0 \u00de \u03c5iII t\u00f0 \u00de 2 66664 3 77775; E\u00bc 1 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 2 6664 3 7775;A\u00bc 0 1 0 0 k m d m 0 0 1 0 1 0 0 1 0 1 2 6664 3 7775;B\u00bc 0 1 0 0 2 6664 3 7775 and the parameters are chosen as: m \u00bc 16; d \u00bc 25; k \u00bc 12"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003019_j.triboint.2017.04.018-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003019_j.triboint.2017.04.018-Figure3-1.png",
        "caption": "Fig. 3. Bearing house with heaters controlled with a PID system.",
        "texts": [
            " Two other thermocouples were used to measure the temperatures of the room and of the air flow around the bearing housing. When assembled in the modified four-ball machine, the rolling bearing assembly was exposed to forced air convection to evacuate the heat generated during bearing operation using two 38 mm diameter fans running at 2000 rpm, cooling the chamber surrounding the bearing housing. In order to control the temperature during the tests, the bearing assembly was mounted with two heaters which were controlled with a PID control system with feedback given by thermocouple III (see Fig. 3). The control system can assure a temperature variation always below than \u22131 \u00b0C. The repeatability of those tests were proved from previous works by developed by Fernandes et al.. The work published in 2013 [21] was repeated in 2015 [22] for the same PAOR lubricant under constant temperature of 80 \u00b0C and an axial load of 7000 N but using different samples and a close internal friction torque measurements with Cylindrical Roller Thrust Bearings were obtained with a maximum difference of \u00b1 18%. This torque variation was accepted since the accuracy of the torque sell is of \u00b1 10%"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001405_spc.2017.8313035-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001405_spc.2017.8313035-Figure2-1.png",
        "caption": "Fig. 2. Coordinate system of robot",
        "texts": [
            " The LQR and H2 controller achieved stabilization for the system but has distortion in the graph when those controllers have been implemented in the hardware [5] which indicates that those controllers are robust, and analysis of control input signal has not been done in this research. Based on the results, the simulation result may not be same with hardware result due to the robustness of the system. The dynamic performance of a balancing robot depends on the efficiency of the control algorithms and the dynamic model of the system. By adopting the coordinate system shown in Fig. 2 using Lagrangian method, it can be shown that the dynamics of the two wheeled EV3 LEGO robot under consideration is governed by the equation (1) to (12) of motion equations. Based on Fig. 2, is the body pitch angle that use gyro value, is the wheel angle that use encoder value and is the yaw of body that use encoder difference. The input motor voltage of both motor is same due to the stabilization system. The parameter of the robot has been tabulated in Table I. Based on Lagrangian method, the Lagrangian equation will formed from the sum of the kinetic energy and potential energy of the system. The Lagrangian equation of the system is shown in [5]. (1) (2) (3) (4) Equations (2), (3) and (4) represents the translational kinetic energy as 1, rotational kinetic energy as 2 which the third term in this equation represents rotation energy of motor armature and the gravitational potential energy of the system as "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003677_1.4042041-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003677_1.4042041-Figure5-1.png",
        "caption": "Fig. 5 Two-dimensional torus and Poincar e section",
        "texts": [
            ", a combination of simple harmonic motions. A simple harmonic motion is equivalent to a one-dimensional sphere S1. Hence, the combination of n asynchronous vibrations and one synchronous vibration is equivalent to the orbit on the surface of (n\u00fe 1)-dimensional torus, i.e., the direct product of (n\u00fe 1) S1 [15] S1 S1 S1 \u00bc Tn\u00fe1 (3) For a simple example where only single RRW and the synchronous motion are present, the motion corresponds to the circulating orbit on the two-dimensional (2D) torus, shown in Fig. 5. Poincar e section is defined as a local plane that intersects with an S1. When the synchronous shaft motion is chosen as the S1, the orbit intersects with Poincar e section at one point, which is called Poincar e point, for one rotation. Thus, a series of the Poincar e points appear on the Poincar e section. A bifurcation diagram is a plot of the signal values at Poincar e points. Since the induction motor had a small slip, which is not exactly known, the actual rotational speed is several Hertz slower than the reference frequency"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002351_s12206-015-0506-2-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002351_s12206-015-0506-2-Figure2-1.png",
        "caption": "Fig. 2. Necessary vectors and frames defined for the wheel(tire model: the inertial(frame (Black), the wheel(frame (Red) and the grid(frame (Blue).",
        "texts": [
            " The stiffness matrix is the tangent stiffness matrix that is obtained by linearizing the high(resolution FE wheel(tire model about an equilibrium state of interest. For reasons of clarity and readability, damp( ing terms are excluded from the equations. It is further as( sumed that all the nodes have only 3 translational DoF, except for the wheel(node. To model the wheel(tire, in addition to defining an inertial : \u2212 \u2212 frame, we define a floating wheel(fixed \u03be \u03b7 \u03b6\u2212 \u2212 frame with its origin at the wheel(node, as shown in Fig. 2. The tire has a symmetric structure about the wheel axis; mean( ing that its mass and stiffness do not change as it rotates. When the tire rotates, new nodes can be defined at the vacant location of the old nodes, using the same mass and stiffness matrices. This suggests that the FE grid could become inde( pendent of the rotation of the tire about its main axis. There( fore, the rotation about the main axis can be described as Eule( rian and the rest of the five DoF of the grid as Lagrangian. In order to formulate such a mixed Eulerian(Lagrangian model, a new grid(frame is defined as \u03be \u03b7 \u03b6\u2212 \u2212 at the wheel, hav( ing a common axis with the wheel(frame, but not rotating about the main axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001927_1.4917498-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001927_1.4917498-Figure2-1.png",
        "caption": "FIG. 2. The removal of hemicellulose and subsequent results. (a) Cellulose, pectin, and hemicellulose are shown in green, blue, and red, respectively. The left image shows the onion epidermal cell layer in dilute sulfuric acid, with a large amount of hemicellulose present. After heating at 120 C for 10 min, most of the hemicellulose is dissolved, as shown in the image in the right. (b) The image in the left is a schematic diagram of the onion epidermal cells plated with gold of different thickness acting as the top and bottom electrodes. The image in the right displays the completed onion actuator. (c) Schematic of the experimental setup and measurement system.",
        "texts": [
            "4% pectin,19 wherein the cellulose and hemicellulose were separated by water. The entanglement between cellulose and hemicellulose fibrils as the cells dry out hinders the cells to be driven in elastic deformation. The hemicellulose is thought to play a role in regulating elongation of the cell wall. To make the dried cells elastic, an acid pretreatment process was utilized to remove the hemicellulose from the cell wall.20,21 After the acid pretreatment, only a small amount of hemicellulose is left in the cell wall of the onion epidermal cells. Fig. 2(a) depicts a comparison of the amount of hemicellulose prior to and after the acid pretreatment process. In order to determine whether or not the hemicellulose had been hydrolyzed and removed from the cell wall interior, we used the phenol-sulfuric acid method to detect the presence of any sugar in our solution.22 So as to accurately represent the amount of sugars in the solution, a standard curve was first prepared23 (supplementary Fig. S1). The treated cell walls were analyzed with x-ray diffraction (XRD)",
            " % at least. To transform the processed onion cells into a functioning actuator, we first sputtered the onion epidermal cell layer with different thicknesses of gold on both sides (top, 24 nm; bottom, 50 nm). The gold layers were intentionally deposited at different thickness as to generate different bending stiffness on the upper and bottom cell walls. This was to make the bending actuation more prominent. Following this, the cell layer was cut into a rectangular strip with dimensions of 25 mm 5 mm (Fig. 2(b)). The long sides of the cells are perpendicular to the length of the electrode23 (supplementary Fig. S3). This is because the measured modulus of elasticity in the perpendicular direction was This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 203.64.11.45 On: Mon, 11 May 2015 18:42:44 smaller than that in the parallel direction23 (supplementary Fig. S4). Thus, a relatively softer and more flexible actuator was obtained. Fig. 2(c) shows the schematic of the experimental setup and measurement system. The free-end deflection of the onion cell actuator was then measured with a laser displacement sensor (KeyenceLK-G3001V). When a voltage from 0\u20131000 V was applied, electrostatic forces deformed the onion cells, causing them to simultaneously bend and, either contract or elongate. Due to the special structure of onion cells, changing the magnitude of the applied voltage will cause the onion cell actuator to bend in different directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002577_fpmc2015-9540-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002577_fpmc2015-9540-Figure6-1.png",
        "caption": "Figure 6 - Fluid mesh of the JB of an EGM",
        "texts": [
            " However, since the solution strongly depends on the geometrical boundaries of the problem, the application of these analytical solutions is unsuitable for cases such as EGMS for which the diameter to width ratio of the bearing approaches the value of 1. A solution for these cases can be obtained by a numerical approach. Since pressure variation along the gap height are neglected in the Reynolds equation, the lubricating interface can be discretized with a 2D mesh. In this work, the Eq. (7) is then solved though a finite volume method, using a preconditioned gradient algorithm with a diagonalized incomplete Cholesky preconditioner [25]. Thanks to the simple geometry of the domain (as shown in Fig. 6) a structured mesh, preferred in finite volume methods, has been used. This allows for a fairly small number of fluid cells. The curvature of the film can be neglected being the gap height much smaller than all the other dimensions of the domain. With that, the fluid mesh can be unwrapped to obtain a more convenient Cartesian reference system, instead of a cylindrical one. The gap height \u2013 the necessary input to solve the fluid pressure according to Eq. (7) \u2013 can be defined with reference to the reference system of Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000851_s12541-015-0332-6-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000851_s12541-015-0332-6-Figure1-1.png",
        "caption": "Fig. 1 CFRP grinding wheel using the carbon fiber lamination method45 (Printed with permission from Ohshita)",
        "texts": [
            " Nonuniform radial deformation occurs in CFRP grinding wheels at ultrahigh speeds. As a result, machining accuracy is affected, along with spallation and damage to the abrasive layer, which seriously affects the safety performance of the grinding wheels. For this reason, Ohshita et al.45 proposed that each layer of carbon fiber sheeting should be rotated by a specific angle around the axis of the wheel. This allowed the sheet material to have a consistent elasticity modulus and coefficient of thermal expansion in the radial direction (Fig. 1). Rotation at ultra-high speeds resulted in uniformly distributed radial centrifugal stress and thermal deformation. Under the same experimental conditions, the machining surface roughness of this new CFRP grinding wheel design was only one-third of that of the traditional CFRP grinding wheel. As a result, it was more suitable for ultra-high speed precision grinding. Although the complexity of the manufacturing process increased, this design scheme gave the CFRP grinding wheel a more obvious tendency towards isotropy, and greatly improved the overall uniformity of grinding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000056_j.precisioneng.2014.11.008-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000056_j.precisioneng.2014.11.008-Figure9-1.png",
        "caption": "Fig. 9. The fabricated test stand.",
        "texts": [
            " Fabrication and testing of a variable preload controlling rototype A prototype of the designed system based on the specifications btained from FE analysis was built. In order to investigate the erformance of the proposed system, it was necessary to meaure the bearing preload variations when the spindle is rotating. hereas in other works, the preload is measured indirectly using he relationship between the preload and spindle stiffness, natual modes and other parameters [16,19,20], in this study, a special etup was designed for direct and real-time measurement of the reload induced by the presented system during spindle rotation. Fig. 9 shows the designed test stand. The prototype was set up n the table of a machining center, and the rotational speed was rovided by machine tool spindle. The spindle did not have an djustable preload. The bearing system, including bearing, shaft, ousing, and the variable preload system, was fixed on a cutting force dynamometer (Kistler model 9257B) which was in turn fixed to the machine table. One end of the shaft was fit in the bearing while its other end was attached to the spindle tool holder"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002210_1.4029828-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002210_1.4029828-Figure1-1.png",
        "caption": "Fig. 1 Differential element of rotating ring supported on elastic foundation, gray shaded area, and under internal pressure",
        "texts": [
            " Results are validated for a uniform homogenous ring using the analytical solution, which was derived for this purpose (see Appendix). The methodology used to derive a spectral element consists of relating the nodal forces and the nodal degrees-of-freedom (displacements and rotations) to the wave amplitudes and then substituting the latter to obtain the relation between forces and degreesof-freedom in the frequency domain, i.e., the dynamic stiffness matrix. A scheme of the rotating curved beam segment considered in this paper is shown in Fig. 1. The centroid of each cross section passes through the centerline. The circumferential coordinate along the centerline is denoted by s and the radial coordinate normal to the centerline is denoted by z. The radial and the tangential centerline displacements are w and u, respectively. The segment may be subjected to an internal pressure p0, and supported by an elastic foundation of radial and tangential stiffness kw and ku, respectively. In this work, the Euler\u2013Bernoulli beam theory is used to model thin rings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003422_s12206-018-0307-5-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003422_s12206-018-0307-5-Figure2-1.png",
        "caption": "Fig. 2. The KOMACHI solid tire series II and size 6.00-9.",
        "texts": [
            " The baggage towing tractor was driven with the speeds between 10 km/h and 50 km/h to tow the baggage carts. The failure or blowout of phenomena might be happened by using the solid tire under the severe conditions such as overloading, high speed, or high temperature work place was out of interest. The solid tires were manufactured by V. S. Industry Tyres Co., Ltd in Thailand which made of the natural rubber. It had been constructed by enveloping with three layers of the natural rubber components. The solid tire brand KOMACHI which is selected to use with the baggage towing tractors is shown in Fig. 2. The pneumatic tire was regular usage with the baggage towing tractors of the Thai Airways International Public Company Limited. It had the outside diameter and width which were approximate as the dimension of KOMACHI solid tire. The pneumatic tire weight was 9.0 kg. Fig. 3 shows the pneumatic tire which was carried out to determine the vibration effects using the drum testing. The specification of KOMACHI solid tires and pneumatic tires are described in Table 1. The width and rim size of the solid and pneumatic tire were 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000776_s11771-012-1135-x-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000776_s11771-012-1135-x-Figure2-1.png",
        "caption": "Fig. 2 Sensor readings to obtain desired virtual bottom path",
        "texts": [
            " Remark 4: High frequency (HF) components of disturbances can cause unnecessary wear and tear of the actuators and must be removed by low-pass filter from the vehicle measurements before they enter the control loop [19]. In the present work, the necessary filtering of HF is assumed to have been taken care of in the output measurements. The INSM tracking controller presented in the previous sections can be applied to bottom-following for AUVs. In this work, a setup is considered where two narrow beam echo sounders, mounted underneath the AUV, scan the seabed along the vehicle\u2019s direction of forward motion, as shown in Fig. 2. As a result of this setup, a preview-based method can be adopted to build the desired virtual bottom path from measurement data of echo sounders [10]. Both the controller design parameters are displayed in Table 3. The environmental current disturbances are considered to validate the performance of bottomfollowing controller for the AUV. The velocity of the current is set as current 0.75, 30 50 0.50, 50 60 0.25, 60 75 0.15, 75 90 z z u z z            J. Cent. South Univ. (2012) 19: 1240\u22121248 1246 And the flowing direction of the currents is 0\u00b0 (joint angle about the negative direction of x-axis)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003135_0954406217720823-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003135_0954406217720823-Figure1-1.png",
        "caption": "Figure 1. Schematic of the ring with a noncircular cross-section on an elastic foundation.",
        "texts": [
            " A model is developed in which the large deformation and the nonlinear terms are considered in the strain, and the model incorporates both the in-plane and out-of-plane bend and the out-of-plane torsion in the displacement fields. The dynamic equations are obtained using the Hamilton variation principle, and the mode expansion is used to obtain the analytical solution for the natural frequency. To verify the accuracy of the ring model, it was applied to the vibration modal analysis of the radial truck tire. The results of the theoretical formulas are proven to be in good agreement with the results of the finite element model and the experimental data. The governing equation of a ring with a noncircular cross-section Figure 1 shows a ring with a noncircular cross-section rotating at a constant speed on an elastic foundation. The inner wall of the ring is under uniform pressure. The elastic properties of the foundation are modeled by distributed springs in the radial, circumferential, and axial directions (ku, kv, and kw, respectively). The equivalent distributed springs are linear elastic. The elastic constants of the strings will be identified by using the FEM results of the modal analysis. The origin of the coordinate system is at the center of the ring"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003770_iros.2018.8594476-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003770_iros.2018.8594476-Figure5-1.png",
        "caption": "Fig. 5. UR3 arm [mm]",
        "texts": [
            " Similarly, sub-actor networks provide the forcing term, f , according to the state which consists of the output of metaactor network, goals, and the output of the canonical system. As the Fig. 4 describes, HRL successfully generates the trajectory to pass through two via-points while normal RL fails and passes only the via-point 1. The following section describes how we applied HRL to the framework of DMPs on the UR3 arm to perform a pickand-place task. The arm platform we used is a 6-DOF robot arm with 1-DOF gripper as in Fig. 5. and the controller available from Universal Robots A/S was used for low-level control of the arm. For experiments, the demonstration includes seven variables, representing the end-effector's position (x,y,z) in the Cartesian space, the end-effector's orientation (q0,q1,q2) in the quaternion space, and the gripper's position (e), which represents \u2018open\u2019 when e = 0 and \u2018close\u2019 when e = 1. Then, we calculate the forcing term through the transformation system in equation (1). We assume that the position of the target is known since perceptual component such as visual processing is currently not the focus of our research"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002908_itsc.2016.7795749-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002908_itsc.2016.7795749-Figure1-1.png",
        "caption": "Fig. 1. Tilt-rotor UAV frames and variables definition.",
        "texts": [
            " The load transportation problem is solved in the presence of parametric uncertainties and constant external disturbances affecting the system. Moreover, constraints on the states and control signals are considered in the optimization problem. This section describes the equations of motion of a Tiltrotor UAV with suspended load obtained through the EulerLagrange formulation. More details about the modeling process can be found in [10]. The complete system can be seen as a multibody system (see Fig. 1): the Tilt-rotor UAV itself composed by three bodies (a main body and two thrusters\u2019 groups) and the suspended load being the fourth body. For modeling purposes, all bodies are assumed rigid and the load is assumed to be attached to the main body by a massless inelastic rod on a spherical joint. Further, two actuators compose each thruster\u2019s group: a servomotor to tilt the propellers and a rotor to generate the lift force. Six frames are defined to describe the Tilt-rotor UAV motion with suspended load: the inertial frame I, and the moving frames B and Ci, which are, respectively, frames rigidly attached to the main body center of rotation and to the center of mass of the i-th body"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003688_042044-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003688_042044-Figure2-1.png",
        "caption": "Figure 2. General view of cylindrical articles on substrates # 0 (a), # 1 (b), # 2 (c) and # 3 (d).",
        "texts": [
            " The I point coordinates will then be (Xi, Yi, Z0) since no displacement was along the OZ axis which will be performed only when forming a new layer. Specifying the trajectory radius R and onset (P.0) one sets up the circular trajectory basis with the numerical control system which then calculates a set of points with required sampling frequency and thus determines the deposition trajectory. There is no need in manual generating the trajectory point sets in order to form the needed shape component. Also wire feed rate and working table displacement velocity are entered to the control system. One can see in figure 2(a) article # 0 made using the electron beam process from SS 321 wire and process parameters shown in table 1. The poor quality of deposition including the unmelted wire segments and rough surface resulted form using too high wire feed rate as well as non-optimal deposition velocity. Articles # 1, # 2 and # 3 in figure 2 have been obtained using the process parameters (table 1) which differed from those of # 0 article by changed deposition velocity. Article # 1 (figure 2b, figure 3) was obtained at 0.1 m/s deposition velocity and showed irregular top edge shape which resulted from excess melting at that low deposition velocity. No unmelted wire segment was observed so that the wire feed rate was at least acceptable value. 3 1234567890 \u2018\u2019\u201c\u201d Freeforming the article # 2 was carried out at 0.23 m/s. The article\u2019s surface is rather smooth without any unmelt wire segments or excess melting the edges (figure 2c). This article may be acceptable from the viewpoint of the layers deposited but has curved walls, probably resulted from their insufficient thickness. Article #3 was obtained at 0.18 m/s (figure 2d, figure 3) and characterized by smooth surfaces with- out any unmelt wire or excess melting. 4 1234567890 \u2018\u2019\u201c\u201d The mechanical characterization of the articles obtained was performed using compression tests on specimens cut from the articles as shown in figure 5. The Testsystems 110M-10 test machine has been used in the process at ambient temperatures. Table 2 contains mean yield stress values for different specimens tested. It follows from the results that article #3 shows the maximum stable mechanical strength distribution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003356_j.apm.2018.01.018-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003356_j.apm.2018.01.018-Figure10-1.png",
        "caption": "Fig. 10. 2D axisymmetric tyre model used for the thermal analysis and the heat transfer boundary conditions (Obtained from: [21] and [15] ).",
        "texts": [
            " This simplification was based on the assumption that heat generation due to friction between the tyre and the road will be negligible compared to that caused by hysteresis effects [3] . In this analysis, the heat transfer boundary conditions applied to the numerical model included the heat generated in the tyre cross-section, the conduction of the generated heat and the convection of the heat from the outer tyre surface to the surrounding environment. The effects of heat transfer by radiation were ignored, since previous studies have shown its effects to be negligible [20] . The 2D axisymmetric tyre model used for the thermal analysis is shown in Fig. 10 . An ambient temperature of 25 \u00b0C was specified for the analysis, and consequently a uniform initial temperature condition of 25 \u00b0C was applied to all of the nodes. The heat generation in the elements was defined as a thermal Volume Flux boundary condition. The flux was calculated from the elastic strain energy density obtained from the rolling analysis using Eq. (3) . The heat generated in the tyre was conducted to the exposed tyre surfaces. The thermal conductivity, specific heat and emissivity of the tyre materials used for the analysis are provided in Table 6 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000413_s12206-011-1201-6-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000413_s12206-011-1201-6-Figure5-1.png",
        "caption": "Fig. 5. Load distribution on roller and bearing.",
        "texts": [
            " (5) The bending moment acting on the tapered rolling bearing can be obtained as 1 2M M M= + (6) where ( )1 0 1 1 1 sin Z n j j i iM L L D q i n \u03b1 = = \u239b \u239e= \u2212 +\u239c \u239f \u239d \u23a0 \u2211\u2211 ( )2 0 2 1 1 sin Z n j j i iM L L D q i n \u03b1 = = \u239b \u239e= \u2212 +\u239c \u239f \u239d \u23a0 \u2211\u2211 where ( )1 jq i and ( )2 jq i can be obtained by Eq. (5). Because there is an implicit function in Eq. (5), bending moment of bearing should be solved using Newton-Raphson iterative method for Eqs. (5) and (6). When the tolerance between ad- jacent calculation results of M is 42 10\u2212\u00d7 , the result can be considered reliable. By the above equations, the bending moment and tilting angle can be obtained. According to the definition of stiffness, the bending stiffness of the tapered roller bearing can be obtained as: r dMk d\u03b8 = . (7) Fig. 5 shows the stress distribution and the moment on sin- gle rolling element and the whole bearing. According to the authors\u2019 past research [18, 19], radial stiffness of tapered roller bearing is ( ) ( ) 0 0 0.13 0.130 1.13 lim 1lim ln ln ( )ln 1 ln 0.13 r r r c F F r r r r r r r r r r r r r FK F F F F F Fn mF m F F C F F n m m F C F \u03b4 \u03b4\u0394 \u2192 \u0394 \u2192 \u2212 \u2212\u0394 \u2192 \u2212 \u0394 = \u0394 = +\u0394 \u2212 \u23a1 \u23a4\u2212 +\u0394 + + +\u0394 + \u23a2 \u23a5\u0394 \u0394\u23a3 \u23a6 = + + + \u22c5 (8) and radial damping of tapered roller bearing is 1 2 1C = 1 1+ C C \u2032 . (9) There is a tilt angle between inner and outer raceways besides radial compression; the tilt angle changes state vectors transfer relation of roller bearing, as Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001084_j.mechmachtheory.2017.05.017-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001084_j.mechmachtheory.2017.05.017-Figure1-1.png",
        "caption": "Fig. 1. The pinion generated by rack-cutter.",
        "texts": [
            " Then, the theory of compensated conjugation is established. Because the tooth modification is based on the transmission function rather than the specific geometry, the new method can be used for gears not only with a high contact ratio, but also with a low contact ratio. Based on the new method, the LTE in helical gears can almost have no variations, which has not been accomplished in most previous literatures. Moreover, after a further longitudinal modification, the PPTE could increase only slightly. As depicted in Fig. 1 , the rack-cutter t 1 generates the pinion. During the generation, the rack-cutter performs a translational motion while the pinion performs a rotational one. In this research, the pinion surface is a modified surface whose deviation from the standard involute tooth surface is controlled by the intentional designed TE and the further longitudinal modification (further details are given below in Section 3 ). Relations between the motion of the rack-cutter and pinion could be observed herein. \u03b81 is the rotational motion of the generated pinion; L 1 is the additional translation parameter of rack-cutter due to the intentional designed TE (for a standard involute gear, L 1 =0); r p 1 is the pitch radius of a pinion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001179_1468087417725221-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001179_1468087417725221-Figure5-1.png",
        "caption": "Figure 5. Real model in the case of valve closed.",
        "texts": [
            " An interesting finding related to case 3 (fully opened valve) depicts a close coincidence between the two cases considering an external flow (solid line) and an internal flow (dashed line), with a difference less than 8W/m2K. Valve closed. The determination of heat transfer from exhaust valve to its seat requires considering the complex geometry and the thermophysical properties of all materials of the cylinder head. The real model of the contact valve-seat is transformed into an ideal model to assess the heat flux conducted from the valve to the seat. This real model (Figure 5) is characterized by the presence of valve, seat, engine block (aluminium) and cooling fluid (water) and supposes that the heat flow path for each medium looks like a truncated cone.8 The equivalent thermal resistance including the conduction within seat and aluminium, the contact resistance between valve-seat and seat aluminium and the convection resistance of water is as follows Rtot =Rt v s + es ksAs +Rt s al + eal kAlAAl + 1 hH2OAH2O \u00f05\u00de where ks and kAl are the conduction coefficient of seat and engine block and hH2O HTC of cooling water"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000056_j.precisioneng.2014.11.008-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000056_j.precisioneng.2014.11.008-Figure8-1.png",
        "caption": "Fig. 8. Contours of maximum stress on the spring and spacer.",
        "texts": [
            "7 162.8 able 2 he calculated preload force in FE model based on rotational speed. Rotational speed (rpm) 1000 2000 3000 4000 5000 6000 7000 Preload force (N) 1445 1410 1333 1219 1055 915 712 parabolic function of rotational speed; i.e. the preload decreases at a higher rate when the rotational speed increases toward higher speeds. 2.4. Stress analysis The finite element analysis is used to find design data for the system prototype. The maximum stress contours on the spring and spacer are shown in Fig. 8. It is observed that stress concentration occurs at the root of the slits because of the sharp edges. The maximum stress in worst case is determined to be as high as 270 MPa. Therefore, a proper material (Mo40 alloy steel) is selected t c 3 3 p o p s W t r s p o p a h In the first step of the test, an initial preload of 1460 N was measured by tightening the nut. Then, the spindle speed was increased gradually up to 7000 rpm in 1000 rpm steps in order to examine o withstand these stresses. Stress concentration may also be overome by drilling small holes at the root of slits"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002679_0309324715614194-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002679_0309324715614194-Figure1-1.png",
        "caption": "Figure 1. Scheme of contact for a (a) separate asperity and (b) wavy surface.",
        "texts": [
            " The comparison between the solutions of these problems makes it possible to study the effect of the complete shape description of rough surface, depending on the sliding velocity, normal load, and adhesive properties of the surfaces. Institute for Problems in Mechanics, Russian Academy of Sciences, Moscow, Russia Corresponding author: Irina Goryacheva, Institute for Problems in Mechanics, Russian Academy of Sciences, pr. Vernadskogo, 101-1, Moscow 119526, Russia. Email: goryache@ipmnet.ru A spherical indenter (1) which is a model of an asperity is loaded by a normal force P and slides with a constant velocity V along a viscoelastic layer (2) in the direction of x-axis as shown in Figure 1(a). The viscoelastic layer lies on the rigid base (3). The shape of the indenter is described by the function f(x, y)= x2 + y2 2R \u00f01\u00de where R is the radius of the asperity tip. The mechanical properties of the viscoelastic layer are modeled by the linear one-dimensional (1D) foundation2 w+Te \u2202w \u2202t = H E p+Ts \u2202p \u2202t \u00f02\u00de where p and w are the pressure and the normal displacement on the boundary of the viscoelastic foundation, E is the elastic modulus, H is the thickness of the viscoelastic layer, and Te and Ts are the retardation and relaxation times, respectively",
            " Thus, the problem is reduced to solving a 2D contact problem for each strip to determine the contact pressure distribution pj(x) in the contact region x2Oj and the boundary of the region of adhesive interaction Oa, after which the friction force can be calculated in accordance with equations (12) and (13). Sliding of a periodic wavy surface Consider a rigid wavy surface sliding with the velocity V along the x-axis on the viscoelastic foundation. The shape of the wavy surface is described by the periodic function f(x, y)= h h 4 cos 2px l +1 cos 2py l +1 \u00f014\u00de where h and l are the height of asperities and distance between them, respectively, h l (Figure 1(b)). Since the wavy surface is periodic with the period l, the contact problem can be considered in a square region x 2 ( l=2; l=2); y 2 ( l=2; l=2). This square contains one asperity of the periodic wavy surface. The conditions of periodicity p(x, y)= p(x+ l, y)= p(x, y+ l) and w(x, y)=w(x+ l, y)=w(x, y+ l) must be satisfied. In the moving system of coordinates (x, y, z), the boundary conditions (7) for the stresses and displacements take place at the foundation surface (z= 0) in the square region x 2 ( l=2; l=2); y 2 ( l=2; l=2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003868_b978-0-12-815889-0.00012-x-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003868_b978-0-12-815889-0.00012-x-Figure1-1.png",
        "caption": "FIG. 1",
        "texts": [
            " Theories by Landau and Mermin, supported by a great many experimental observations connoted that 2D crystals are thermodynamically unstable and cannot exist in free state (8, 9). Despite all these strong theoretical and experimental, the 2D allotrope of carbon was found to exist and was first isolated through mechanical exfoliation in 2004 by Geim and coworkers (9). GP is the first proof that 2D materials can exist in stable crystalline form. It can be considered as the fundamental element for not only 3D graphite but also 1D CNTs and 0D fullerenes (10). Fig. 1 shows how 2D GP is related with graphite, CNT, and fullerene. GP, graphite, CNT, and fullerenes. 2513 PROPERTIES OF GRAPHENE Immediately after its discovery, GP caught the attention of researchers all over the world for its unique physical and chemical properties and has proved its theoretical importance beyond any doubt. As a result of its outstanding electric properties, GP has been utilized in fabricating various ECSs, biosensors, and high-power supercapacitors (10). In this book chapter, we will mainly focus on electronic properties of GP and related materials and their application in fabricating ECSs for medicinal applications"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001088_s11432-016-9025-9-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001088_s11432-016-9025-9-Figure3-1.png",
        "caption": "Figure 3 Scheduled path.",
        "texts": [
            " When the schedule for the kv-th UAV is completed, if there are still remaining sets of the path to be connected, proceed to schedule the path for the (kv + 1)-th UAV. Otherwise, the whole schedule is over. Suppose that there are s sets of the path scheduled for the kv-th UAV. Then, pth,kv is given as follows: pth,kv = ( Rtg,h1i1k1 \u00b7 \u00b7 \u00b7 Rtg,hrirkr srtg,rr+1 Rtg,hr+1ir+1kr+1 \u00b7 \u00b7 \u00b7 Rtg,hsisks ) , srtg,rr+1 = ( ep,hrirkr sp,hr+1ir+1kr+1 ) . (18) The scheduled path to search the rectangles is shown in Figure 3. The scheduled path needs to be connected with the take-off and landing locations for the assigned UAVs. The objective of the connection is to avoid collision. There are two ideas. One is that there is no joint between any two channels and one UAV takes one channel. The other is that more than one UAVs share one channel and there is no meeting chance for the UAVs in the channel. The first idea needs sufficient space to allocate the channels for the UAVs. The second idea needs time allocation for the UAVs to share a channel",
            " The complete path for the three UAVs is scheduled as follows: path,k = ( ptof,k pth,k plnd,k ) , k \u2208 {k1, k2, k3}, ptof,k = ( tof,1k \u00b7 \u00b7 \u00b7 tof,4k sgi,k ) , k \u2208 {k1, k2}, ptof,k3 = ( tof,1k3 \u00b7 \u00b7 \u00b7 tof,4k3 ) , plnd,k = ( sgo,k lnd,4k \u00b7 \u00b7 \u00b7 lnd,1k ) , k \u2208 {k1, k2, k3}, (22) where as the additional segments are added, the related segments are revised as follows: tof,3k = ( t2,k t\u03043,k ) , tof,4k = ( t\u03043,k t\u03044,k ) , k \u2208 {k1, k2}, t\u03043,k1 = ( t2x,k1 ci \u2212 dsfty/2 ) , t\u03044,k1 = ( psx,k1 ci \u2212 dsfty/2 ) , t\u03043,k2 = ( t2x,k2 ci + dsfty/2 ) , t\u03044,k2 = ( psx,k2 ci + dsfty/2 ) , lnd,3k = ( l\u03043,k l2,k ) , lnd,4k = ( l\u03044,k l\u03043,k ) , k \u2208 {k1, k2, k3}, l\u03044,k1 = ( pex,k1 co \u2212 dsfty ) , l\u03043,k1 = ( l2x,k1 co \u2212 dsfty ) , l\u03044,k2 = ( pex,k2 co ) , l\u03043,k2 = ( l2x,k2 co ) , l\u03044,k3 = ( pex,k3 co + dsfty ) , l\u03043,k3 = ( l2x,k3 co + dsfty ) . The complete scheduled path is shown in Figure 3. XnOnYn denotes the North-East-Down (NED) frame. 1 and 2 denote the two take-off and landing locations in which the local region is enlarged for display. This completes the development of the online RSA. A 2D grid map is drawn with the missed areas to verify the developed online RSA. The results of the four steps are shown in Figure 4. The computed results show that the developed online RSA is successful to decide the number of the UAVs to be assigned and to schedule the complete path for the assigned UAVs to take off to search the missed areas and fly back to land"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure2-1.png",
        "caption": "Fig. 2. (a) Description of the overlap between the particles (b) Kelvin\u2013Voigt model with Coulomb\u2019s friction.",
        "texts": [
            " In 2D simulations, the gyroscopic forces are not taken into account. By using the smooth Discrete Element Method (DEM), initially developed by Cundall and Strack [27,28] , the contact forces in a bearing are described with a model depending on an elastic force displacement law, Coulomb\u2019s friction and viscous damping. The contact occurs only when discrete elements with radii a i and a j interpenetrate, which means that a contact between a rolling element and a ring or a contact between a rolling element and a cage component exists as suggested by the Fig. 2 (a). The equivalent model of contact is described with a Kelvin\u2013Voigt model, Fig. 2 (b) considering an analogy with a mass-spring-damper system. The dry friction coefficient \u03bc = 0 . 3 is overestimated for numerical reasons. According to the Stribeck curve, an important value of friction coefficient describes a boundary lubrication regime [12] . For important loads and low rotation velocities, the lubricant film can not be formed and therefore is not taken into account. In bearing applications, a typical friction coefficient varies from \u03bc = 0 . 001 to \u03bc = 0 . 1 , according to the lubricant regime relative to operational conditions",
            " This kind of description reserves the possibility of introducing more realistic viscous models in dynamics also called Elastohydrodynamic lubrication theory [12,29] (EHD), assuming the lubrication regime may be estimated or determined [16,30,31] . The force F i between particles at the interface includes the inter-particle interaction forces and the external forces. F i = F ext i + \u2211 j = i F j\u2192 i (1) where F ext i are the external forces on particle i (gravity, loading, ...). F j\u2192 i is the force exerted by particle j on particle i. F j \u2192 i is deduced from analogies with a damp-spring mass system. From Fig. 2 (a), this model includes a normal and a tangential component. F j\u2192 i is decomposed as follows: F j\u2192 i = F n n + F t t (2) F n is the contact force in the normal direction and F t is the contact force in the tangential direction. By introducing an analogy with a damped spring mass system, the coefficients K n , C n and K t , C t are introduced, representing the stiffness and the viscous damping of contact, in the normal direction n and in the tangential direction t respectively, and the overlap between particles u = u n n + u t t estimates the contact force: { F n = K n \u00d7 u n + C n \u00d7 \u02d9 u "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001211_978-981-10-6250-6-Figure2.11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001211_978-981-10-6250-6-Figure2.11-1.png",
        "caption": "Fig. 2.11 The Internet-based test rig",
        "texts": [
            " VFG scheme, stable, \u03c4\u0304sc = 1, \u03c4\u0304ca = 1 0 50 100 150 200 250 300 350 \u22120.1 \u22120.05 0 0.05 0.1 0.15 Time x 1(t) Example 2.3 The system matrices for the system in (2.1) are set as A = ( 0.7 0.2 0.3 0.5 ) , B = ( 0.05 0.2 ) ,C = ( 1 0 ) . This system can be shown using Theorem 2.2 to be stable under \u03c4\u0304sc = 3, \u03c4\u0304ca = 2, Nu = 8, Np = 10. The simulation result is illustrated in Fig. 2.10. Example 2.4 In this example, an Internet-based test rig is used to verify the effectiveness of the packet-based control approach. This test rig consists of a plant (DC servo system, see Fig. 2.11a) which is located in the University of Glamorgan, Pontypridd, UK, and a remote controller which is located in the Institute of Automation, Chinese Academy of Sciences, Beijing, China (see Fig. 2.11b). The plant and the controller are connected via the Internet, whose IP addresses are Fig. 2.9 Example 2.2. FFG scheme, unstable, \u03c4\u0304sc = 1, \u03c4\u0304ca = 1 0 50 100 150 200 250 300 \u22121.5 \u22121 \u22120.5 0 0.5 1 1.5 x 106 Time x 1(t) Fig. 2.10 Example 2.3. Packet-based control, stable, \u03c4\u0304sc = 3, \u03c4\u0304ca = 2 0 50 100 150 200 250 \u22120.25 \u22120.2 \u22120.15 \u22120.1 \u22120.05 0 0.05 0.1 Time x 1(t) The DC servo system is identified by [118] to be a third-order system and in state-space description has the following system matrices, A = \u239b \u239d 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000782_j.ifacol.2015.06.433-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000782_j.ifacol.2015.06.433-Figure2-1.png",
        "caption": "Fig. 2. Additive Manufacturing process",
        "texts": [
            " According to [8], some stages of the process by addition of materials have been determined:  Modeling for obtaining the 3D model on CAD system;  CAD models conversion to \"STL\" format (surface tessellation language or standard tessellation language);  Check if no error occurred in the conversion;  Creation of fixtures;  Orientation to manufacturing (vertical / horizontal);  Slicing and preparaion for manufacturing (construction parameters and numerical control program);  Manufacturing (model execution in machine);  Post-processing (removal of fixtures and resin excess, post-curing and finishing of surface). Additive Manufacturing enables the fast, flexible and costefficient production of parts directly from 3D CAD data \u2013 a technology that helps you to perform your tasks in an innovative way [9]. INCOM 2015 May 11-13, 2015. Ottawa, Canada 2393 2320 Silva, R. J. et al. / IFAC-PapersOnLine 48-3 (2015) 2318\u20132322 Figure 2 illustrates the steps of adding layers process from the design to the final part (end item): Due to its relative operational simplicity, welding has become the most important industrial process used for manufacturing of metal parts. Considered as a joining method, this process or its variations are used for the deposition of material on a surface in order to recover worn parts or to form a coating with special characteristics. It was from the 19 th century that the welding technology has emerged on the world scenario contributing to the apperance of fusion welding processes with the discovery of acetylene by Edmund Davy, Sir Humphrey Davy experiences (1801- 1806) with the electrical arc and with the development of production sources of electricity [10]",
            " Develop a controlled mechanical device named torch holder, to generate speed and constant or variable position on displacement axis for deposition welding layers. 2. Check parameters such as: level of current, current pulse frequency, operating voltage, arc length, torch travel speed in translation of axes, feed speed of filler metal, composition of filler material, distance from the outlet point to the part, torch angle related to the piece, type of INCOM 2015 May 11-13, 2015. Ottawa, Canada Silva, R. J. et al. / IFAC-PapersOnLine 48-3 (2015) 2318\u20132322 2321 Figure 2 illustrates the steps of adding layers process from the design to the final part (end item): Fig. 2. Additive Manufacturing process 2.3 Welding Due to its relative operational simplicity, welding has become the most important industrial process used for manufacturing of metal parts. Considered as a joining method, this process or its variations are used for the deposition of material on a surface in order to recover worn parts or to form a coating with special characteristics. It was from the 19 th century that the welding technology has emerged on the world scenario contributing to the apperance of fusion welding processes with the discovery of acetylene by Edmund Davy, Sir Humphrey Davy experiences (1801- 1806) with the electrical arc and with the development of production sources of electricity [10]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002380_j.ifacol.2015.08.056-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002380_j.ifacol.2015.08.056-Figure7-1.png",
        "caption": "Figure 7: Landing scenario components.",
        "texts": [
            " The magnetometer measurements are used to calculate the yaw of the UAV and the difference between this value and the current estimation is used as the measurement of the yaw error. In a similar way, the outputs of the accelerometers are used to calculate the measurements of the roll and pitch errors. This block is in charge of calculating the relative position of the landing point to the helicopter. This is done through the measurement of the tether system, the altimeter and the results of the AHRS. In Figure 7 a graphical description of the landing scenario with all the components that take part in the conversion block is shown. The first step in this block is to obtain the relative position measurement in the contact point (CP) of the tether with the fuselage. The CP of the vehicle and the altimeter are not placed at exactly the same position. This spatial separation causes small differences in position and velocity due to the lever-arm effect. In this landing phase it is necessary to have as much accuracy as possible; the measurement of the altimeter must be corrected according to equations (1) and (2): (1) (2) Once the altitude to the contact point has been calculated, it is possible to obtain the position relative to the landing point in the navigation axes",
            " The magnetometer measurements are used to calculate the yaw of the UAV and the difference between this value and the current estimation is used as the measurement of the yaw error. In a similar way, the outputs of the accelerometers are used to calculate the measurements of the roll and pitch errors. Conversion This block is in charge of calculating the relative position of the landing point to the helicopter. This is done through the measurement of the tether system, the altimeter and the results of the AHRS. In Figure 7 a graphical description of the landing scenario with all the components that take part in the conversion block is shown. Figure 7: Landing scenario components. The first step in this block is to obtain the relative position measurement in the contact point (CP) of the tether with the fuselage. The CP of the vehicle and the altimeter are not placed at exactly the same position. This spatial separation causes small differences in position and velocity due to the lever-arm effect. In this landing phase it is necessary to have as much accuracy as possible; the measurement of the altimeter must be corrected according to equations (1) and (2): (1) (2) Once the altitude to the contact point has been calculated, it is possible to obtain the position relative to the landing point in the navigation axes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001165_icuas.2017.7991362-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001165_icuas.2017.7991362-Figure2-1.png",
        "caption": "Fig. 2. Unmanned Aerial-Underwater Vehicle coordinate frames.",
        "texts": [
            " Given an unit axis of rotation k\u0302 and a rotation angle \u03b1 , the vector part of the quaternion of rotation q\u0304 and the scalar part q0 satisfy q0 = cos(\u03b1/2), q\u0304 = k\u0302sin(\u03b1/2) (4) Rotation of a vector by a quaternion q can be represented by means of a rotation matrix R(q) of the form R(q) = I3 +2q0q\u0304\u00d7+2q\u03042 \u00d7 (5) where I3 stands for the three by three identity matrix and q\u0304\u00d7 is the skew-symmetric matrix of the vector q\u0304. Unit quaternions present ambiguities since they double cover the SO(3) group (R(q) = R(\u2212q)). For further details about quaternion algebra refer to [16]. UAUVs can be modeled as a rigid body evolving in a three-dimensional space, where the medium properties may change abruptly, i.e. density, viscosity. Let us consider an earth-fixed inertial coordinate frame I and a body fixed coordinate frame B (see Figure 2). Define the position vector \u03be = [x, y, z]T , along with the quaternion of rotation q describing the vehicle\u2019s orientation. Then, using the NewtonEuler formalism, the equations of motion of an air-water vehicle UAUV (refer to [17], [18] for more details about modeling of air, respectively underwater vehicles) can be found as m\u03be\u0308 = T R(q)e3 \u2212g(m\u2212\u03c1V )e3 \u2212\u03c1D\u03be (\u03be\u0307 ) (6) J\u2126\u0307 = \u03c4 \u2212\u2126\u00d7J\u2126\u2212\u03c1D\u2126(\u2126) (7) with a mass m, gravity constant g, volume of the vehicle V and inertia matrix J. For simplicity, we consider a perfect alignment of the centers of mass and buoyancy with the body-fixed frame B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure6-1.png",
        "caption": "Fig. 6. Influence of center distance error on the contact region.",
        "texts": [
            " Thereinto, the linear error is considered as the center distance error \u03c0x = 0.03, the tangential error \u03c0y = 0.03 and the axial error \u03c0z = 0.03, and the angle error is considered as the perpendicularity error \u03b5x = 0.03 of hole series and the angle error \u03b5y = 0.03 of Y2-axis. In order to explore the influence of different error sources and provide the basis for the tolerance determination, other errors are omitted when the influence of one error is discussed. (1) Influence of center distance error Fig. 6 shows the deflection tendency of contact region in TL (tooth-length) and TH (tooth-height) directions of pinion and gear when the center distance error \u03c0x = 0.03. Here, the red zone represents the range of total contact region which is formed by overlapping instantaneous contact regions, the gold curve represents actual contact locus (CL) of centers of instantaneous contact regions and the turquoise curve represents theoretical CL of contact points. Fig. 6(b,d) presents the boundary of 2D contact region, consisted of five instantaneous contact regions, and the deflection values of actual CL in TL and TH directions. It can be found that:(a) relative to the theoretical state, when the center distance error is positive, total contact regions will deflect to the topland of pinion and the root of gear when the tooth pair conjugates from the toe to the heel, respectively; (b) the influence of center distance error on the contact region will gradually reduce in the conjugation process of teeth, and the shift of bearing contact tends to zero in TL direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000599_j.engfailanal.2012.02.008-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000599_j.engfailanal.2012.02.008-Figure2-1.png",
        "caption": "Fig. 2. Traverse of casting pedestal.",
        "texts": [
            " In the technological flow it serves for transportation of ladles with liquid steel between converters and tundish of machine for continuous casting. During the operation, the structure of casting pedestal is exposed to dynamic loading due to transported mass of ladles with liquid steel, because the pedestal carries out on roller bed reversible rotational movement around vertical axis ov (by 180 ) and the traverse of pedestal tilts around horizontal axis oh (by approximately 7.0 ) (Fig. 1). The welded carrying system of casting pedestal consists of middle part and two beams (Fig. 2). The traverse beams are connected with middle part by bolt and wedge joints. After more than 10 years of operation of casting pedestal arose the problems with releasing of bolt and wedge joints. Subsequently, there were found out cracks in the supporting structure of pedestal that lead to proposals of modifications of the structure. In the paper are given results obtained during assessment of crack initiation causes in the supporting structure of pedestal as well as the measures for ensuring its further operation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000500_s11665-012-0227-y-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000500_s11665-012-0227-y-Figure5-1.png",
        "caption": "Fig. 5 The 3D-FE bending model of thin-walled rectangular tube",
        "texts": [
            " Establishment of the Springback Prediction Model Considering the Variable Young\u2019s Modulus for the Rotary-Draw Bending of Thin-Walled Rectangular 3A21 Tube The whole process of a bending tube includes three stages: the bending process, the retracting mandrel process, and the springback process. Since the bending process and the retracting mandrel process can be calculated by a bending model, only two different FE models are needed to be established in each process. An FE model for the rotary-draw bending process and the retracting mandrel process based on the literature (Ref 18) has been established (as shown in Fig. 5). Some key problems were resolved in the original model, such as geometry modeling, assembly modeling, boundary condition, load definition, etc. However, the variation of Young\u2019s modulus versus plastic strain was not considered. Hence, to improve the simulation accuracy of the bending process, the constitutive model built in section 2-4 is implemented into the FE model by VUMAT under the platform of ABAQUS/Explicit. Since springback involves no contact and only mild nonlinearities, the output files of the results in the bending process are imported into ABAQUS/Standard software, such as the final shape of the rectangular tube, related information of elements, nodes, the stress, strain, etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002340_1350650114562485-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002340_1350650114562485-Figure3-1.png",
        "caption": "Figure 3. Calculating the contact strain in a bearing.",
        "texts": [
            " In order to solve it, the authors used a procedure used before,2,8 consisting of iterative searching of such dislocations in bearings (shifts and tilts), for which all conditions of equilibrium are fulfilled: \u2013 compliance of radial reactions of bearings with radial forces influencing the bearings in accordance with the scheme in Figure 2, \u2013 compliance of the sum of the axial reactions of bearings with the sum of the external axial loads, \u2013 compliance of angles of bearings tilts and angles of deflection of shaft under bearings, \u2013 compliance of bending moment in bearings and bending moments considered when calculating the line of shaft deflection and \u2013 compliance of the sum of axial shifts in bearings with the value of initial tightness. Calculating the reaction forces and moments appearing in bearings is an important element in this draft procedure. The calculations are based on assuming dislocations of the inner ring in relation to the outer one, i.e. shift in relation to three axes (fx, fy, fz), as well as tilts in relation to two axes (yy i yz). The dislocations lead to contact deformation at the contact of each ball with the race, calculated according to the relation from (1) to (7), and shown in Figure 3. The section presented in Figure 3 is located in any plane (between x\u2013y plane and x\u2013z plane) w \u00bc rbw \u00fe rbz 0, 5 dbz dbw\u00f0 \u00de \u00f01\u00de PQ \u00bc rbw \u00fe rbz Dk \u00f02\u00de AQ \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0PQ\u00de2 w2 q \u00f03\u00de BQ \u00bc AQ\u00fe fx rp y sinc\u00fe z cosc \u00f04\u00de BP0 \u00bc w\u00fe fy cosc\u00fe fz sinc \u00f05\u00de P0Q \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi BP0\u00f0 \u00de 2 \u00fe BQ\u00f0 \u00de 2 q \u00f06\u00de d \u00bc P0Q PQ \u00f07\u00de In the relations presented above, c stands for the angle of a ball position calculated from the plane (x\u2013y)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002435_iros.2015.7353998-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002435_iros.2015.7353998-Figure4-1.png",
        "caption": "Fig. 4. Modified bicycle model of a bevel tip needle. The frames {C} and {B} are attached to front and back wheels. The back wheel is located at needle tip",
        "texts": [
            " Omni-directional wheels shown in Figure 3 can move independently in two orthogonal directions. Such wheels satisfy the wheel plane constraint (3) but violate the non-slip constraint (4). Orthogonal to the wheel plane, the motion constraint will be Bvy \u2212 \u03c9R = 0 (11) in which \u03c9R denotes the rotation velocity of the rollers. This is the degree of freedom added to the system allowing the wheel to have lateral movements. Here, we replace the back wheel of the bicycle with the omni-directional wheel of Figure 3, as shown in Figure 4. In this Figure, \u03b2, ` and \u03b1 denote the fixed front wheel angle, the distance between the two wheels and the rotation angle of the needle tip in body frame {B}, respectively. This rotation of the needle tip from the insertion direction by angle \u03b1 is due to the lateral movement of the back wheel causing the needle to be tangent to the non-circular path. The inputs u1 and u2 denote the insertion velocity along the z axis of frame {B} (which equals Bvz) and the rotation velocity of the needle about its axis, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002813_s00202-016-0441-y-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002813_s00202-016-0441-y-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of a ball bearing",
        "texts": [
            " It can be predicted from the bearing characteristic geometry and the rotation speed of themechanical shaft [23]. \u23a7 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 fo(Hz) = fr Nb 2 ( 1 \u2212 Db Dp cos\u03a6 ) fi (Hz) = fr Nb 2 ( 1 + Db Dp cos\u03a6 ) fc(Hz) = fr 1 2 ( 1 \u2212 Db Dp cos\u03a6 ) fb(Hz) = fr Dp Db ( 1 \u2212 D2 b D2 p cos2\u03a6 ) (1) where \u2013 fb the characteristic frequency of a ball, \u2013 fi the characteristic frequency of the inner ring, \u2013 fo the characteristic frequency of the outer ring, \u2013 fc the characteristic frequency of the cage, \u2013 fr the rotation frequency of the mechanical shaft, \u2013 Nb the number of balls, \u2013 Db the ball diameter (see Fig. 1), \u2013 Dp the pitch diameter (see Fig. 1), \u2013 \u03a6 the contact angle (see Fig. 1). We consider in this work that the bearings are radial, i.e., with a contact angle\u03a6 = 0\u25e6.Moreover, these frequencies are calculated with the consideration that the contacts ball/ring are perfectly punctual and that the balls are rolling without sliding. In real bearing, the balls slide at the same time as they roll on the slopes. To take into account this sliding phenomenon, a multiplicative sliding factor of frequency is introduced, defined as the ratio between the rolling distance and the slidingdistance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000072_aim.2014.6878331-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000072_aim.2014.6878331-Figure1-1.png",
        "caption": "Fig. 1. Workpiece",
        "texts": [
            "klimchik at mines-nantes.fr path. Section 3 describes the robotic cell. Section 4 deals with the formulation of a mono-objective optimization problem to find the optimum placement of the workpiece with regard to a proposed machining quality criterion. Section 5 highlights the optimum and worst workpiece placements within the robotic cell. The best and worst redundancy planning schemes associated with those placements are also determined. Section 6 is about the conclusions of the paper and the future work. Figure 1 illustrates the workpiece to be machined, which is made up of aluminum alloy. FW of origin OW is the frame attached to the workpiece. The five segments AB, BC, COW , OWD and DE, of length equal to 200 mm each, have to be milled. The tool path is offset by the tool radius from the five segments to be milled. The tool path is discretized into n points and is shown in dashed line in Fig. 1. Frame FPi is attached to the ith point of the tool path, i = 1, . . . , n. XPi is along the feed direction. ZPi is along the tool axis and points toward the robot. The machining quality is affected by the robot deviation due to the cutting forces applied on the tool [2]. The cutting 978-1-4799-5736-1/14/$31.00 \u00a92014 IEEE 1716 conditions are given in Tab. I where fz , ap and ae denote the feed rate, the depth of cut and the width of cut, respectively. The cutting forces are evaluated thanks to the cutting force model described in [3]",
            " Closed loop chain and frames F0, F6, F7 and FW \u03b2i Spindle Z7 XPi Fig. 5. Kinematic redundancy characterized with angle \u03b2i 0TW is the homogeneous transformation matrix from frame F0 to frame FW expressed as: 0TW =  2Q2 1 \u2212 1 \u22122Q1Q4 0 0xOW Q1Q4 2Q2 1 \u2212 1 0 0yOW 0 0 1 0zOW 0 0 0 1  (8) 0xOW , 0yOW and 0zOW being the Cartesian coordinates of point OW expressed in frame F0. WTPi is the homogeneous transformation matrix from frame FW to frame FPi attached to the ith point of the tool path as shown in Fig. 1. 0T6 is the homogeneous transformation matrix from frame F0 to frame F6. 6T7 is the homogeneous transformation matrix from frame F6 to frame F7 and depends on the geometry of the spindle and how the latter is mounted on the robot end-effector. Here, 6T7 =  0 \u2212 \u221a 2/2 \u221a 2/2 0 1 0 0 0 0 \u221a 2/2 \u221a 2/2 0.684 0 0 0 1  (9) 7TPi is the homogeneous transformation matrix from frame FPi to frame F7. Note that the KUKA KR270-2 robot has six degrees of freedom, whereas the milling operation sets only five degrees of freedom as the rotation of the spindle about the tool axis is not fixed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003603_1350650118800581-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003603_1350650118800581-Figure1-1.png",
        "caption": "Figure 1. Schematic view and CAD model of floating-ring bearing: (a) schematic view of a floating-ring bearing; (b) plain cylindrical floating-ring; (c) a circumferential groove in the outer surface of the floating-ring; (d) four axial grooves in the inner surface of the floating-ring; and (e) a circumferential groove in the outer surface and four axial grooves in the inner surface of the floating-ring.",
        "texts": [
            " Although the high evaluation capacity of modern computers and the most advanced solvers can be used, the numerical solutions of the Reynolds equation for both inner and outer oil-films performed with numerical methods, such as finite difference method (FDM) and finite element method (FEM), are considered slower in comparison to direct formulas for the bearing\u2019s impedance forces yielded from non-numerical solutions such as short bearing approximation,1,2,11\u201314 exact analytical solution,15\u201318 and other approximate analytical model.19,20 However, the incorporation of holes or axial/circumferential grooves in the floating ring leaves no room for non-numerical treatment of the Reynolds equation. The pressure and shear stress distributions in the oil-films vary due to the modified boundary conditions introduced by the holes or the axial/circumferential grooves. The bearings of the floating-rings usually contain grooves to maintain oil supply (see Figure 1(a)). Lubricant flowing from the outer film to the inner film plays an important role in maintaining an uninterrupted oil-film and removing most of the frictional heat to cool the inner film of the floating-ring. The land areas reduce when the full-floating-rings featuring grooves, and the load capacities and bearing torques of the oil-films would reduce as well, exerting impact on the nonlinear oscillations of TC rotors. However, literature in this field rarely concern axial or circumferential grooves in full-floating-ring bearings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003234_1350650117738395-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003234_1350650117738395-Figure1-1.png",
        "caption": "Figure 1. Scania\u2019s current (2017) single cone synchronizer.",
        "texts": [
            " Sliding friction, sliding wear, boundary lubrication, synchronizer, gear shift, temperature, -comp Date received: 24 February 2017; accepted: 12 September 2017 Due to the different gear ratios in the gearbox, there will be a rotational speed difference between the engaging bodies when a gear shift is initiated. Synchronizers are used to synchronize the rotational speed of the gearbox during gear shifts in manual transmissions (MT), dual clutch transmissions (DCT), and automated manual transmissions (AMT). An example of a synchronizer is shown in Figure 1. The gear wheel to be engaged is free to rotate around the shaft it is mounted on. The coupling disc is mounted on the gear wheel, and the inner cone is mounted on the coupling disc. The driver component is mounted on the shaft, and the latch cone, shift sleeve and wire spring are mounted on the driver. The driver, and therefore the shift sleeve, is rotationally locked to the shaft. When a gear is engaged, the shift sleeve locks the rotation of the gear wheel to the shaft via the driver. The shift sleeve presses the latch cone against the inner cone and the resulting torque seeks to synchronize the rotational speed of the gear wheel and the shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002547_1.4033101-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002547_1.4033101-Figure6-1.png",
        "caption": "Fig. 6 Coupled rotor\u2013cartridge model",
        "texts": [
            " The reaction forces from the SFDs are calculated with the corresponding moments being taken about the outer ring center of mass. The SFDs act in parallel; therefore, the reactions are summed. The squeeze film Reynolds number is given as Re \u00bc qxC2 l (12) where q is fluid mass density. Inertial forces are considered significant when Re > 1 [24]. The current system has a small radial clearance resulting in minimal added mass contribution with Re 1. Therefore, fluid inertia effects were neglected. 3.3 Rotor\u2013Cartridge Model. Figure 6 illustrates the coupled rotor\u2013cartridge model. The cylindrical rotor has modified properties making it dynamically equivalent to the actual turbocharger rotor. The material properties consist of an aluminum compressor, steel shaft, and Inconel turbine. The density of a portion of the compressor and turbine was increased in order to represent the additional mass of the blades. The equivalent rotor was developed in lieu of a more geometrically representative model due to the number of elements that would be required"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001848_j.procir.2015.03.095-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001848_j.procir.2015.03.095-Figure2-1.png",
        "caption": "Fig. 2. Simple 3D dexel model of a gear shape constructed by I disks and J ray points per each disk",
        "texts": [
            " Due to complexity of the single chip simulations performed in previous work [10,11], the direct enhancement to the full hob is numerically inefficient due to expanding of calculation time \u2013 the single simulations by setup 2 took by favorable discretization about 40 hours for only one generating position. Comparing to that, a new model for the geometric description has to finish the simulation in reasonable time. Hence, a dexel approach is proposed for this purpose. The classical dexel [12] structure represents only the surface of the body as two dimensional set of scalar values (ray points, see Fig.2): The polar coordinates with discretization along angle are well suitable for the disk geometry description (Fig.2-3). The full gear is assembled in a set of the dexel disks, discretized with the proper quality, as well as for height (i-index) and for the radial shape (j-index). The reasonable discretization for the gear (module pitch 2.7 mm) is 1 degree (J=360) for each disk and 1 mm of height, which results in I=27 disks and 360 nodes for simplest dexel structure. If a higher resolution is needed, the dexel model can be resolved properly fine, time cost can be held almost linear proportional due to parallelization of disks"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002547_1.4033101-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002547_1.4033101-Figure3-1.png",
        "caption": "Fig. 3 Cartridge cross section",
        "texts": [
            " In order to predict the dynamics of the turbocharger system, a coupled rotor\u2013cartridge model was developed. A description of a shaft supported by two separate deep groove ball bearings is given by Brouwer et al. [15]. The preceding model has been extended to represent the turbocharger rotor\u2013cartridge system that is under consideration. Details of the model are described in Secs. 3.1, 3.2, and 3.3 below. 3.1 Cartridge Design. A cross section of the angular contact ball bearing cartridge is depicted in Fig. 3. The cartridge contains a single outer ring with outer raceways on opposing ends in a back to back configuration. The inner rings are split in the middle to allow assembly. The use of a common outer ring requires modification of the contact routine between a ball and outer raceway. The centroid of the outer raceway is not located at the center of mass of the outer ring. An additional vector is needed to locate the outer raceway centroid with respect to the outer ring center of mass. The position of each outer raceway centroid is defined with respect to the center of mass of the cartridge outer ring by vectors in the body fixed reference frame as shown in Fig",
            " Therefore, the stiffness of the bearing was estimated with the equations given by Gargiulo [21]. According to these methods, a value of 0.3 N s/mm was approximated and applied to the inner raceways as described by Brouwer et al. [15]. 3.2 SFD. Additional bearing damping is not necessary for low speed applications, but at high speeds, damping is essential as large displacements may occur while traversing critical speeds. For this reason, the rotor cartridge is surrounded by SFDs. Two SFDs are located on each half of the cartridge as shown in Fig. 3. Oil is supplied to the groove between each pair of SFDs. End seals are not utilized in the current configuration. The SFDs are modeled using the short bearing formulation. The nonlinear reaction forces from short SFDs are presented by Gunter et al. [22] and Taylor and Kumar [23]. With a circular orbit and assuming a p film, the normal and tangential forces simplify to FN \u00bc 2lRL3xe2 C2 1 e2\u00f0 \u00de2 (10) FT \u00bc plRL3xe 2C2 1 e2\u00f0 \u00de 3 2= (11) where l is the fluid dynamic viscosity, x is the orbital velocity, e is the eccentricity ratio, R is the radius with a value of 14"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003379_0954410018764472-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003379_0954410018764472-Figure2-1.png",
        "caption": "Figure 2. Separation diagram of the multi-rigid bodies.",
        "texts": [
            " To simplify the modeling process, we suppose that the left wing and the right wing will realize the synchronous motion in the morphing process and the symmetry axis of the fuselage lies in the same plane as the wings on both sides. By comparison with the conventional modeling approach for the morphing aircraft, which treats it as a single rigid body, the Kane modeling can describe the aircraft dynamic behavior more accurately by taking the whole structural changes of the each moving part into consideration (Figure 2). Thereby, the dynamic of the morphing motion is described by using Kane method in this paper. We suppose that the morphing aircraft consists of five separate rigid bodies: fuselage(Body1), inner part of left wing(Body2), outer part of left wing(Body3), inner part of right wing(Body4), and outer part of right wing(Body5), and their masses are mb,m1,m2, m1,m2, respectively. The important parameters of the morphing aircraft are given in Appendix 1. Firstly, the appropriate coordinate systems must be created in order to set up the Kane equations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001463_j.matpr.2018.03.041-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001463_j.matpr.2018.03.041-Figure2-1.png",
        "caption": "Fig 2. Finite Element Analysis of plate without hole.",
        "texts": [
            " Jeyaprakash.P. / Materials Today: Proceedings 5 (2018) 14526\u201314530 14529 The plate normal is aligned in z direction and plate area is located on xy plan. The simply supported boundary conditions are applied on each four edges by fixing the translational displacement in z axis. The compression load is applied uniformly along the two opposite edges. These plates are meshed with quadratic composite shell elements based on first order shear deformation theory. The buckling of composite plates is shown in fig. 2, 3(a), (b), and(c) During the buckling analysis of plate without hole in the ANSYS conducted for the following bidirectional laminate. The buckling loads for cross-ply laminate composite plates with the cutout diameter to the laminate with (a/b) varying from 0.0 to 7.0 are given in Table 4 the result of are displayed in Fig 3 (a), (b) and (c). The buckling load is increased by changing the position of cut out from top to bottom. It is worth mentioning that, for the cut out range of a/b=0.0 to a/b=7, the buckling load have a considerable change compare to perfect plate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002331_1350650114559997-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002331_1350650114559997-Figure5-1.png",
        "caption": "Figure 5. Projected definition of fundamental spall area.",
        "texts": [
            "comDownloaded from inner or outer race will deteriorate first, the pitch diameter is used in this calculation. Typically, the inner race will show signs of deterioration first due to the reduced contact area (leading to higher local stress for a given load compared to the outer race). The inner race radius could be used; however, it would have minimal impact on the final limit. The arc length is determined using the following equations: SP \u00bc DP 2 \u00f017\u00de \u00bc 360 n \u00f018\u00de DP \u00bc do \u00fe di\u00f0 \u00de 2 \u00f019\u00de Step 4: Determination of fundamental defect area. Figure 5 shows the composite projected defect area known as the fundamental defect area (AFD). This area is determined using the following equation by adding: 1. the contact area, 2. the area formed by the pitch sector and the contact area width, and 3. the area formed by the transverse spall growth. AFD \u00bc AC \u00fe SP Dy \u00fe 2 y \u00f020\u00de Step 5: Determination of reliability factor. Once the fundamental defect area has been determined, it is then modified by applying a reliability factor. The reliability factor is used to modify the allowable spall area based on the operating context of the subject machinery"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001375_jas.2017.7510895-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001375_jas.2017.7510895-Figure5-1.png",
        "caption": "Fig. 5. Peltier actuated part.",
        "texts": [],
        "surrounding_texts": [
            "In the above results, there is still a fundamental issue, that is, by using the isomorphism-based factorizing method, we can get two factors N and D\u22121, while in the proof of the coprimeness of RCF, AN + BD = M , the form and value of D is necessary. However, if D\u22121 is highly nonlinear or too complex, then the verification work will become too difficult. Therefore, a sufficient condition is proposed by which the right factorization can be proved to be coprime which means that the NFS is stable.\nTheorem 1: Let us consider the well-posed NFS which is demonstrated in Fig. 1, if B\u22121 is designed to linearize the factor D\u22121 such that,\nMD\u22121B\u22121 \u2212APB\u22121 = I (10)\nwhere the given operator M \u2208 \u00a7(U ,Y ), then the right factorization is coprime, i.e. the NFS is overall stable.\nProof: From Fig. 1, it can be seen that r(t) = (I + AND\u22121B\u22121)(e)(t), moreover, according to the notation of well-posedness, e(t) is uniquely determined by r(t), thus I + AND\u22121B\u22121 is invertible\n(I+AND\u22121B\u22121)\u22121\n= [(BD + AN)D\u22121B\u22121]\u22121 = BD(BD + AN)\u22121 (11)\nbased on (9) and the assumption of M , we find that\n(I + APB\u22121)\u22121 = (MD\u22121B\u22121)\u22121 = BDM\u22121 (12)\nthen according to (11) and (12)\nAN + BD = M (13)\nwhich is verified to be a Bezout identity, then N and D\u22121 is the RCF of P , with the result in [7], which also means that the NFS is overall stable. \u00a5\nTherefore, by the proposed sufficient condition (10), there is no need to do the heavy work of obtaining the factor D, which is more economical for real practices. Moreover, based on the idea of LN, the sufficient condition can be extended into the UNFS shown in Fig. 3.\nB. Improved Robust Condition\nThe proposed condition (10) is the complementary and developed design scheme of satisfying the Bezout identity by using the operator-based RCF method, how to extend the\nsufficient condition to the UNFS is the next concern in this paper. Similarly based on LN, the robust conditions can also be developed as follows.\nTheorem 2: Let us consider the well-posed UNFS which is demonstrated in Fig. 3, where P = ND\u22121, P + \u2206P = (N+\u2206N)D\u22121 which respectively denotes the factorization of the nominal and the perturbed plant, G and F are respectively assumed to be G = MD\u22121B\u22121 and F = (A(P +\u2206P )B\u22121\u2212 APB\u22121), if B\u22121 linearizes D\u22121 while (6) and (10) hold, such that\nM\u0303D\u22121B\u22121 \u2212A(P + \u2206P )B\u22121 = I (14) \u2016FG\u22121\u2016 < 1 (15)\nthen the UNFS is overall stable. Proof: Since D\u22121 is linearized by B\u22121, then\n\u2016FG\u22121\u2016 = \u2016[A(P + \u2206P )B\u22121 \u2212APB\u22121]BDM\u22121\u2016 = \u2016[A(P + \u2206P )D \u2212APD]M\u22121\u2016 = \u2016[A(N + \u2206N)\u2212AN ]M\u22121\u2016 < 1 (16)\nthen (9) is satisfied. Moreover, according to Theorem 1, (14) implies that (8) holds, while (10) means that (7) is satisfied. Therefore, since (6)\u2212(9) are all satisfied, then according to Lemma 6, the whole UNFS is robustly stable. \u00a5\nTherefore, if D\u22121 is too highly nonlinear such that it is difficult to obtain its inverse D, then we can design and control the UNFS according to Theorem 1 and Theorem 2. Moreover, the conditions (6), (10), (14) and (15) are denoted to be the improved robust conditions which are feasible for more general cases.\nThe main contribution of the results is that the two sufficient conditions are not just complementing the condition (1) and Lemma 1, they are more developed and can act as a substitute method. In details are as follows:\n1) For the nominal NFS, the condition (1) is extended into (10) under the case of complex or highly nonlinear D\u22121.\n2) For the UNFS, the robust conditions in Lemma 2 [8] ((6)\u2212(9)) are extended to the improved robust conditions ((6), (10), (14), (15)).\n3) The derived two conditions are more suitable for the real dynamics since they consider the relationships among the reference input, feedback and the error signal, instead of using the quasi-state signal which is hard to be obtained [7], [8].\nFor the purpose of demonstrating the validity of the proposed control schemes, a robust control on the temperature of a thermal process will be discussed. It consists of a Peltier actuated device, aluminum plate, temperature sensors, a serial communication and computers which are shown in Figs. 4\u22126. Peltier devices have a particular effect of a pyroelectric element and a so-called Peltier effect in which one side is endothermic while the other side is exothermic when the current flows, moreover, when the direction of the current is reversed, the cold side and the hot side exchange. The",
            "Peltier effect denotes the presence of heating or cooling at the electrified junction of the different conductors. Heat will be generated or removed at the junction whenever the current is made to flow through it.\nBased on corresponding physics rules, such as Fouriers law concerning thermal conduction, specific heat capacity equation, electrothermal amount by Peltier effect, thermal conduction by temperature difference (of course, all the heat generated cannot be decided by the Peltier effect alone, as it is also affected by Joule heating, thermal gradient effects\nand other rules), a mathematical model is set up where the parameters are demonstrated as shown in Table 1.\nwhere \u03b15t1i denotes the heat which flows from the endothermic part to the radiation part owing to the Peltier effect, \u03b14(ts \u2212 t1) shows the heat movement produced by the temperature difference of the two parts of the device, for simplification, we choose t0 \u2212 tx to be process output y, thus the mathematical model of the Peltier-actuated NFS is re-expressed as\ny(t) = 1\n\u03b11m e\u2212Qt\n\u222b eQ\u03c4 ( \u03b15t1i(\u03c4)\n\u2212 \u03b14(ts \u2212 t1)\u2212 1 2 Rpi(\u03c4)2\n) d\u03c4\nQ = \u03b12(2S1 + 2S2 \u2212 S3) + 2\u03b13S4 d1\n\u03b11m . (17)\nt1 =(l1 + l7)i + l2e \u2212l3t \u2212 l8e \u2212l9t\n(ts \u2212 t1) = (l4 + l10)i + l5e \u2212l6t \u2212 l11e \u2212l12t (18)\nFrom the physical nature and the mathematical model, we find that the temperature varies with i as well as i2, thus the control system derived from the temperature of the thermal process is typically nonlinear. Moreover, from the rule of mathematics and the real practice, the current should be considered to be the control input. However, in most of the existing results, the Joule exothermic heat is usually taken to be the control input instead of the current. In this simulation, the robust control will be discussed under the assumption of the current i being the control input, that is, we will deal with the temperature control by using the proposed method based on the operator theory and the corresponding variables are outlined in Table II [11].\nWe assume that the nominal as well as the perturbed plant is respectively right factorized into P + \u2206P = (N + \u2206N)D\u22121,",
            "(P + \u2206P )(u)(t) = 1\n\u03b11m e\u2212Qt\n\u00d7 \u222b eQ\u03c4 ( \u03b15t1u(\u03c4)\u2212 \u03b14(ts \u2212 t1)\u2212 1\n2 Rpu\n2(\u03c4) ) d\u03c4 + \u03b4\nD\u22121(u)(t) = \u22121 2 Rpu 2(t) + \u03b15t1u(t)\u2212 \u03b14(ts \u2212 t1) N(w)(t) = 1\n\u03b11m e\u2212Qt\n\u222b eQ\u03c4w(\u03c4)d\u03c4\n(N + \u2206N)(w)(t) = 1\n\u03b11m e\u2212Qt\n\u222b eQ\u03c4w(\u03c4)d\u03c4 + \u03b4. (19)\nThe obtained right factor D\u22121 is seen to be nonlinear which seems to be difficult to acquire the operator D, thus, as proved in Theorem 3.2, the stable controllers are respectively designed\nA(y)(t) = (\u03b11m\u2212 \u03b21) (\ndy(t) dt\n+ Qy(t) )\nB(u)(t) = \u03b22ud(t) (20)\nwhere\nud = \u22121 2 Rpu 2(t) + \u03b15t1u(t)\u2212 \u03b14(ts \u2212 t1) (21)\nthen for simplification, we choose\nM(w)(t) = \u03b23w(t) (22)\nwhere \u03b2i is the gain for adjustment, if\n\u03b23 + \u03b21\n\u03b11m = \u03b22 + 1 (23)\nholds, the sufficient conditions are satisfied:\nMD\u22121B\u22121(e)(t)\u2212APB\u22121(e)(t)\n= [(M \u2212AN)D\u22121B\u22121](e)(t) = [(\u03b23 \u2212 \u03b11m\u2212 \u03b21\n\u03b11m ) 1 \u03b22 ](e)(t)\n= I(e)(t) (24)\nthus (10) is satisfied, then nominal NFS is stable by designed controllers.\nNext, we verify the inequality condition as follows, if the uncertainty\n\u03b4 < \u03b23\nQ(\u03b11m\u2212 \u03b21) (25)\nthen (15) is satisfied as follows:\n\u2016FG\u22121\u2016 = \u2016[A(N + \u2206N)\u2212AN ]M\u22121\u2016 = \u2225\u2225\u2225\u2225 (\u03b11m\u2212 \u03b21)Q\u03b4\n\u03b23 \u2225\u2225\u2225\u2225 < 1. (26)\nTherefore, according to the Theorem 2, (6), (10), (14) and (15) are all satisfied, then the robustness of the UNFS is guaranteed.\nThe simulation results are demonstrated in Figs. 7\u22129 where the quasi-state signal as well as the plant output are both BIBO stable with the parameters \u03b21 = 0.5m\u03b11, \u03b22 = 0.5, \u03b23 = 1, \u03b4 = 0.1.\nTwo main results are obtained in this paper, wherein a sufficient condition in the form of equality is derived to deal with the BIBO stability of the NFS since there exists a case where it is difficult to get the form and value of D because D\u22121 is highly nonlinear or too complex. Based on the LN, the sufficient condition is extended to UNFS and the developed robust conditions are derived as confirmed by the simulation results of robust control for the temperature of a thermal process. In such a case, the proposed sufficient conditions are not just complements but developments of the control and design of the UNFS.\nThe authors would like to thank the group\u2019s work in Tokyo University of Agriculture and Technology."
        ]
    },
    {
        "image_filename": "designv11_13_0000172_roman.2013.6628527-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000172_roman.2013.6628527-Figure5-1.png",
        "caption": "Fig. 5. Frames of captured video depicting the 3 block stacking task",
        "texts": [
            " Using the initial and final images of the green and blue block combination, the initial and final markings that were extracted are \u00b5initial = (1, 1, 0, 0, 0) and \u00b5final = (0, 0, 0, 1, 1). The path found via the Petri net generation algorithm is t1t2 which is the expected sequence of transitions to perform the desired task. Thus, the Petri net is able to generalize to new object combinations. B. 3 Block Setting In this test setting, three primary colored blocks were stacked by a human demonstrator who placed two of the blocks side by side in the middle of the workspace first, before placing the third block on top of the first two blocks as shown in Figure 5. For this test setting the Petri net has to be able to model a combination of a parallel (concurrent) and a sequential task, i.e. the order of actions for placing the first two blocks can be interchanged, but the action of placing the third block must come last. The two test cases considered for this setting are exact imitation and multiple demonstrations. Subset imitation and minor generalization capabilities were also verified. As before, the results of each test case (including the subset imitation and minor generalization) were used to let the NAO robot imitate the task in simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000964_140982283-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000964_140982283-Figure1-1.png",
        "caption": "Fig. 1. An example of the \u03b3-strict tangent cone.",
        "texts": [
            " In other words, a supporting hyperrectangle of a bounded set C is an axis-aligned minimum bounding hyperrectangle. Definition 5. Let A \u2282 R d be an axis-aligned hyperrectangle and \u03b3 > 0 a constant. The \u03b3\u2212strict tangent cone to A at x \u2208 R d is the set (3.2) T\u03b3(x,A) = { cs(A) if x \u2208 ri(A), T (x,A) \u22c2 k\u2208I{z \u2208 R d : |\u3008z,\u2212\u2192rk\u3009| \u2265 \u03b3Dk(A)} otherwise, where I = {k \u2208 D : x \u2208 rbk(A)}, rbk(A) denotes the two facets of A perpendicular to the axis \u2212\u2192rk , and Dk(A) = max(A)k \u2212min(A)k denotes the side length parallel to the axis \u2212\u2192rk . Figure 1 gives an example of the \u03b3-strict tangent cone to A at x. D ow nl oa de d 11 /3 0/ 15 to 1 32 .2 03 .2 27 .6 1. R ed is tr ib ut io n su bj ec t t o SI A M li ce ns e or c op yr ig ht ; s ee h ttp :// w w w .s ia m .o rg /jo ur na ls /o js a. ph p Copyright \u00a9 by SIAM. Unauthorized reproduction of this article is prohibited. Fig. 2. Convex hull, supporting hyperrectangle, strict tangent cone, and feasible vectors f ip satisfying Assumption 3. 3.4. Agreement metrics. We next define uniformly asymptotic agreement and exponential agreement in this section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002541_j.jsv.2016.01.019-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002541_j.jsv.2016.01.019-Figure1-1.png",
        "caption": "Fig. 1. Principle of the experiment. (For interpretation of the references to color in this figure caption, the reader is referred to the web version of this paper.)",
        "texts": [
            " Section 2 describes the tribometer, the optical viewing system, and the elastomer sample. Section 3 presents the experimental results. The characteristics of squeal noise are analysed and the modal analysis of the sample is performed. Measurements of the squeal noise level and velocity weakening friction are also presented. Section 4 is a discussion on the role of the evolution of contact area during the squeal. Finally, some conclusions are drawn in Section 5. The principle of the experiment is shown in Fig. 1. It consists of observing the contact during the occurrence of squeal. An elastomer sample is pressed against a glass disk with a fixed normal load. The contact is lubricated by water. The disk rotation speed is maintained constant during recording of the noise. The radius of the disk is 45 mm and the mean distance between the sample and the centre of the disk is 22.5 mm. For all the experiments reported in this paper, the contact is lubricated by 10 ml of distilled water introduced by a pipette near the contact zone",
            " The second decoder VD06 operates by Doppler effect and delivers the vibrational velocity with a resolution of 0.05 \u03bcm s 1. The signals corresponding to the normal force, tangential force, vibrational velocity and vibrational displacement are acquired synchronously by an analogue-to-digital data acquisition card at the sampling frequency 20 kHz. In order to perform interferometric image of the contact, the glass disk is made of quartz and is coated with a chromium layer which is itself covered by a silica layer of thickness 10 nm (Fig. 1). The sample is therefore in contact with silica. The incident light beam is divided into two beams, one is reflected by the chromium layer, and the other one goes through the silica layer and the interface before being reflected by the elastomer surface. The difference of optical path between the two beams generates an interferometric image where the pixel colour depends on the distance between the chromium layer and the elastomer. A camera with a resolution of 2448 2050 pixels is equipped with a varying zoom allowing a full size image from 7 mm 5",
            " According to the Hertzian theory of non-adhesive elastic contact, the theoretical contact width Dtheo is defined as, Dtheo \u00bc 2 ffiffiffiffiffiffiffiffiffiffi 4FR \u03c0lE r (2) where F\u00bc1 N is the normal load, R \u00bc 0:5 mm the reduced radius of curvature, l\u00bc30 mm the contact length and E the reduced Young's modulus at 23 \u00b0C and 1 Hz defined as 1 E \u00bc 1 \u03bd2elast Eelast \u00fe 1 \u03bd2glass Eglass \" # (3) with \u03bdelast \u00bc 0:5, \u03bdglass \u00bc 0:2, Eelast \u00bc 9:3 MPa, Eglass \u00bc 69 GPa. Numerical application gives Dtheo \u00bc 97 \u03bcm. The theoretical width is inferior to the measured one, indicating that adhesive effects may occur during the elastomer\u2013glass friction in lubricated condition. The contact is heterogeneous and composed of spots. A spot is defined as a set of contiguous pixels having the same colour. The colour of each spot is related to the optical path \u03b4 as shown in Fig. 1. One may define two kinds of spots. The \u2018contact spots\u2019 colour corresponds to where the optical path \u03b4 reaches its minimal value. Indeed the optical path is minimal when the distance between the sample and the disk is minimal. The other spots are called \u2018non-contact spots\u2019. These spots can have different colours corresponding to various separation distances between the elastomer and glass. Thus, there is only one colour defining the contact spots and several colours for the non-contact spots. In order to identify the contact spots, a first experiment is performed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000123_j.snb.2015.03.092-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000123_j.snb.2015.03.092-Figure6-1.png",
        "caption": "Figure 6. Synthesis of mestranol.",
        "texts": [
            " Based on the data interpretation and Mestranol was prepared by Colton et al. (8) as f o l l o w s : Estrone {I} i s converted to its 3-methoxy analog (11) by reaction with methyl sulfate. The ethynyl group may then be introduced at position 1 7 either through reaction with sodium acetylide in liquid ammonia followed by hydrolysis of the sodoxy compound, or through Grignardization with ethynyl magnesium bromide. Almost the sole product of the ethynylation reaction is that which results from attack of reagent from the least hindered a-side of the steroid,Fig.6. 3 86 HUMEIDA A. EL-OBEID AND ABDULLAH A. AL-BADR MESTRANOL 387 4. Absorption, Metabolism and Excretion Mestranol is a synthetic estrogen which is more potent than estradiol. It is readily absorbed from the gastrointestinal tract and is slowly metabolized and excreted in urine and feces. The absorption, metabolism and excretion of the drug has been extensively studied in animals and humans I Fig. 7. Wijmenga and Van der Molen (9) reported a biological halflife of mestranol of 50 hours and that a small proportion of the drug was excreted in milk of nursing mothers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003669_0954406218811869-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003669_0954406218811869-Figure2-1.png",
        "caption": "Figure 2. Analysis of forces imposed on the ball.",
        "texts": [
            "2 \u00f08\u00de Suppose the inner raceway rotates and the outer raceway is fixed, the ratio of the ball rotation speed and the cage speed can be formulated as !R ! j \u00bc 1 cos o,j\u00fetan sin o,j 1\u00fe 0 cos o,j \u00fe cos i,j\u00fetan sin i,j 1 0 cos i,j 0 cos \u00f09\u00de The ratio of the ball revolution speed and the cage speed can be formulated as !m ! j \u00bc 1 0 cos i,j 1\u00fe cos i,j o,j \u00f010\u00de where the attitude angle of the j-th ball can be written as follows tan j \u00bc sin o,j cos o,j \u00fe 0 \u00f011\u00de where 0 \u00bc D dm \u00f012\u00de By conducting the force balance analysis, equations of the j-th rolling ball depicted in Figure 2 can be available as follows Qij sin ij Qoj sin oj Mgj D lij cos ij loj cos oj \u00bc 0 \u00f013\u00de Qij cos ij Qoj cos oj \u00fe Mgj D lij sin ij loj sin oj \u00fe Fcj \u00bc 0 \u00f014\u00de where the contact force between the j-th ball and raceways can be, respectively, written as Qij \u00bc Kij ij 1:5 and Qoj \u00bc Koj oj 1:5. The parameter K means the Hertz contact deform- ation coefficient.12 To judge the contact state between the j-th ball and inner raceway, the following formulations can be available L1 \u00bc BD sin 0 \u00fe z \u00fe Ri x cos \u2019j y sin\u2019j \u00f015\u00de L2 \u00bc BD cos 0 \u00fe x sin \u2019j \u00fe y cos \u2019j fo 0:5\u00f0 \u00deD\u00fe oj \u00f016\u00de L \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L1 2 \u00fe L2 2 q fi 0:5\u00f0 \u00deD \u00f017\u00de When the value of L is less than zero, which means that the ball has gone out of the inner raceway, and the contact state between the j-th ball and inner raceway is detached",
            " Suppose the contact angle between the ball and outer raceway is just equal to zero; hence, the following force balance equation can be written as Qoj Fcj \u00bc 0 \u00f018\u00de where the contact angles between the j-th rolling ball and raceways can be formulated as follows cos oj \u00bc X2j fo 0:5\u00f0 \u00deD\u00fe oj \u00f019\u00de sin oj \u00bc X1j fo 0:5\u00f0 \u00deD\u00fe oj \u00f020\u00de cos ij \u00bc A2j X2j fi 0:5\u00f0 \u00deD\u00fe ij \u00f021\u00de sin ij \u00bc A1j X1j fi 0:5\u00f0 \u00deD\u00fe ij \u00f022\u00de By conducting the force balance analysis, the load balance equations of the inner raceway depicted in Figure 2 can be given as follows Fx Xk j\u00bc1 Qij cos ij \u00fe lijMgj D sin ij sin j \u00bc 0 \u00f023\u00de Fy Xk j\u00bc1 Qij cos ij \u00fe lijMgj D sin ij cos j \u00bc 0 \u00f024\u00de Fz Xk j\u00bc1 Qij sin ij lijMgj D cos ij \u00bc 0 \u00f025\u00de Mx Xk j\u00bc1 Qij sin ij lijMgj D cos ij Ri \u00fe lijMgjfi cos j \u00bc 0 \u00f026\u00de My Xk j\u00bc1 Qij sin ij \u00fe lijMgj D cos ij Ri \u00fe lijMgjfi sin j \u00bc 0 \u00f027\u00de where the position angle of the j-th ball is j \u00bc 2 Z j 1\u00f0 \u00de \u00fe !t \u00f028\u00de where ! means the cage-rotating speed, which can be deduced as follows ! \u00bc 1 2 1 D dm cos 0 !rotor \u00f029\u00de Hence, we can infer that the period of the ball passage is Tvc \u00bc 2 =Z "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure14-1.png",
        "caption": "Fig. 14. FEM models of planetary gear with carrier.",
        "texts": [
            " The corresponding comparative result is therefore not presented. Because double tooth pair contact occurs in the case of annulus-planet gear pairs, the shared load and the contact stress must be checked. As the diagrams in Figs. 12 and 13 illustrate, the analysis has a good converged results when the unit number is also greater than 225. With consideration of the calculation efficiency and accuracy, the unit number is selected as 529 in the following calculation. The FEM model of the planetary gear set, illustrated in Fig. 14 , consists of the carrier with the planet shafts, Fig. 14 (b), and the gears, Fig. 14 (c) and (d). The influences of the twist of the sun gear shaft and the deformation of the carrier are involved in the FEM analysis. The element type is Hex8 (6 faces and 8 nodes). No special elements are necessary to be placed at the points of contact for solving the contact problem by using the FEM software (MSC.Marc), because the special algorithm, so-called Solver Constrain Method, is applied. The FE-model includes 577,970 elements and 715,730 nodes. The setting of the boundary conditions is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002554_1.4033128-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002554_1.4033128-Figure10-1.png",
        "caption": "Fig. 10 Incorporation of dimples at inlet side and outlet side: (a) inlet incorporation of dimples and (b) outlet incorporation of dimples",
        "texts": [
            "org/about-asme/terms-of-use point, the convergence ratio decreases significantly as Hd increases further, The behavior of the convergence ratio in the case of outlet incorporation is identical to that in the case of inlet incorporation. However, with outlet incorporation, the pad is more inclined than for inlet incorporation at any Hd. The maximum K of 0.334 is observed at Hd\u00bc 0.9. As shown previously, it was found that the incorporation of multiple dimples at either side of the lubricated area develops the entire wedge action over the area. On the other hand, when multiple dimples are created in the central zone, the pad is not stabilized. Figure 10 shows a schematic of the multiple-dimples arrangements in the cases of inlet incorporation and outlet incorporation to investigate the influence of the dimple portion width. Figure 11 compares the load W, friction F, and convergence ratio K for various numbers of dimples N between the cases of inlet incorporation and outlet incorporation. When the number of dimples N is 1, a single dimple is incorporated at the side of the lubricated area. This is shown as the \u201cfirst dimple\u201d in Fig. 10. In the multiple-dimples case, dimples are arranged from the side of the lubricated area. In Fig. 10, dimensional specifications ldw and llw for each dimple are the same. Only the number of dimples, N, changes to increase the dimple portion width ld. For comparison between inlet incorporation and outlet incorporation, almost the same load can be obtained in both the cases, although the trends are different and depend on the number of dimples. In the case of the inlet incorporation, the load W increases as the number of dimples N increases. At N\u00bc 6, the maximum value of the load W is located. As N increases further, the load W turns to decrease to obtain no equilibrium solution of moment",
            " The outlet incorporation of dimples could also generate a positive pressure distribution over the lubricated area in the same manner as the inlet incorporation. This trend is important for lubricated areas with reciprocating motion. Incorporation of dimples at either side of the lubricated area may be effective during all of the reciprocating motion period because both incorporation methods could improve the load-carrying capacity and reduce friction. The incorporation of dimples at the center of the lubricated area resulted in no balance of moment. This trend can be observed in Fig. 10, in which no balance solution can be found as the dimple portion width expands from the sides of the lubricated area toward the center. Another important trend was the superiority of the step film shape of the Rayleigh step bearing compared with that created by multiple-dimples incorporation. This fact has also been pointed out in the parallel case [30\u201332]. According to results shown in Figs. 13 and 14, when the dimple located at the inlet side was opened to the inlet, the load capacity was more improved than when all dimples were enclosed within the lubricated area"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000219_016918611x607699-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000219_016918611x607699-Figure1-1.png",
        "caption": "Figure 1. Rimless wheel model.",
        "texts": [
            " In Section 3, we formulate a steady two-period gait based on two recurrence formulas of kinetic energy just before impact. In Section 4, the validity of the analysis is investigated through numerical simulations of VPDW. Finally, Section 5 concludes the paper and describes future research directions. D ow nl oa de d by [ U ni ve rs ity o f W es t F lo ri da ] at 0 9: 34 3 0 D ec em be r 20 14 F. Asano / Advanced Robotics 26 (2012) 155\u2013176 157 We first describe the stability mechanism of a rimless wheel. Since the detailed theory was already explained in Ref. [18], we only outline it here. Figure 1 shows the model of a rimless wheel. Let \u03b1 (rad) be the angle between the frames and \u03b8 (rad) be the angle with respect to vertical. We assume that the total mass, M (kg), is concentrated at the central point and the leg frames have no mass. We also assume that 0 < \u03b1 < \u03c0/2. Let K\u2212 (J) be the kinetic energy just before impact; it satisfies the following recurrence formula: K\u2212[i + 1] = \u03b5K\u2212[i] + E, (1) where i is the step number, \u03b5 (\u2014) is the energy-loss coefficient and E (J) is the restored mechanical energy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003620_j.matpr.2018.06.241-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003620_j.matpr.2018.06.241-Figure1-1.png",
        "caption": "Fig. 1 Bicycle model with dimensions",
        "texts": [
            " These rigid bodies are force transmitting members and are fixed to front fork and rear axle. The seat post and the steering tube were also simulated as rigid links. A similar study was also performed by Pazare et al [3] , where the frame was considered as a truss to find the stresses theoretically in each member and then to compare it with FE results using commercial package Ansys APDL. The tubes were assumed as line i.e. beams elements and subjected to loads as mentioned in previous literatures as shown in below Fig 1 and Fig 2. The approach involves in validating two existing research papers [1] and [2]. The former approach involves just the frame where the loads are applied directly to the frame members. The loads being taken from the references [4] and [5]. The different tubes are named and their respective diameters and length are also given. A uniform thickness of 1mm is given to all the tubes and radius of 4mm is maintained at all the junction of the tubes. Whereas, in the latter approach the loads are applied at the rigid links to simulate real world conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003202_j.ijsolstr.2017.10.008-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003202_j.ijsolstr.2017.10.008-Figure3-1.png",
        "caption": "Fig. 3. Segments of equilibrium path for Dome I under p 1 .",
        "texts": [
            " Different sets of guesses may lead to different equilibrium configurations. Segments of the equilibrium path are followed manually. For the case of force control, the stability of an equilibrium state is determined by applying the Mathematica subroutine FindMinimum to see if V has a local minimum there. For loading at the crown, the equilibrium path is plotted as nondimensional force p 1 versus nondimensional downward crown displacement \u03b41 with p 1 \u2265 0 and \u03b41 \u2265 0. Segments of interest are presented in Fig. 3 for the ranges 0 \u2264 \u03b41 \u2264 2.5 and 0 \u2264 p 1 \u2264 100. Solid curves correspond to Green\u2013Lagrange strain and dashed curves to engineering strain. The dashed curves are indistinguishable from the solid curves except in a neighborhood of point E. h a t r b t c r s W t d m F j t I r p m b c b 2 m E T t ( i p w a a t a p s n p r d d p E a r a t p o d i u 3 s t a m 0 a j g F a F 3 p m c b t F m d eight of joints 10 and 11 is at the left, then the heights of movble joints 5, 1, and 2, and the zero height of joints 8 and 13 is at he right. Equilibrium configurations to which the dome snaps are epresented by dashed lines. All shapes depicted in this paper are ased on analyses using Green\u2013Lagrange strain. Under unilateral displacement control in Fig. 3 , the dome iniially follows path A-B-C. At C, the associated force p 1 has dereased to zero, and joint 1 is at the same level 6.332 as the ing joints 2\u20137 ( Fig. 4 ). This configuration was previously decribed in Lu et al. (2009); Rezaiee-Pajand and Estiri (2016c) , and atson et al. (1983) . (The second reference reported the height of he movable joints as 6.328.) Then the interior of the dome snaps ownward to the centrally inverted configuration D, at which the embers are unstrained. As the indentor pushes further, the dome follows path D-B \u2032 -Ein Fig. 3 . At F, again p 1 has decreased to zero, and now the ring oints are at the same level as the support joints ( Fig. 4 ). The dome hen snaps downward to the unstrained, totally inverted shape G. f the indentor continues to push further downward, the equilibium path follows the segment from G through E \u2032 to larger dis- lacements. It is noted that the maximum compressive strain is approxiately 1.5% at E and F ( Tables 2 and 3 ), so that, in practice, mem- er buckling and/or nonlinearly elastic or plastic behavior may ocur before point F is reached, in which case the analysis would not e valid for the larger values of \u03b41 in Fig. 3 . .4. Force control In this section the downward force p 1 on joint 1 is increased onotonically. Equilibrium shapes corresponding to points A, B, B \u2032 , , and E \u2032 in Fig. 3 are depicted in Fig. 5 , and data are listed in able 3 . As p 1 is increased from zero, the equilibrium path reaches he limit point (local maximum) B at p 1 = 3.1558 with h 1 = 7.447 assuming Green\u2013Lagrange strain), the same values as computed n Koohestani (2013) ; in Rezaiee-Pajand and Estiri (2016a ), the re- orted values are p 1 = 3.1559 and h 1 = 7.445. The dome snaps to B \u2032 , ith snap-through of the section inside the ring ( Fig. 5 ). With further increase of p 1 , the behavior depends on the strain ssumption. Consider Green\u2013Lagrange strain first. At p 1 = 86.076, bifurcation point occurs at E. (For comparison, using interpolaion on data sent to the author by Paolo Valvo gives p 1 = 86.101 t the bifurcation point.). Fig. 6 depicts the neighborhood of this oint, where the dashed curve represents the bifurcating path (not hown in Fig. 3 ). On that path, the equilibrium configurations are ot symmetric about the z axis, and the ring is wavy (e.g., at 1 = 86.0: h 1 = \u20131.880, ring joints 2, 4, and 6 have height 3.524, and ing joints 3, 5, and 7 have height 3.650). As one moves downward on the bifurcating path from E, \u03b41 ecreases (albeit slightly). If \u03b41 had increased instead, then uner unilateral displacement control ( Section 2.3 ) it would be exected that the dome would follow the bifurcating path from point rather than segment D-F in Fig. 3 ( Plaut 2015a,b ). Under force control, equilibrium states on the bifurcating path nd the solid path from E to F in Fig. 3 are unstable. When p 1 eaches 86.076 at E, the dome snaps to point E \u2032 , for which all movble joints are below the support plane ( Table 3 ; Fig. 5 ). If engineering strain is assumed, no bifurcation point is obained in Fig. 3 (dashed curves). In Table 3 , the second line for oint E gives data at the limit point at p 1 = 88.655, and the secnd line for point E \u2032 lists data for the configuration to which the ome snaps from the limit point. For high strains, the analysis usng Green\u2013Lagrange strain is more appropriate than the analysis sing engineering strain ( Blandford, 1996a ). . Dome I loaded at a ring joint In this section, Dome I is loaded at joint 2, with nondimenional force p 2 and nondimensional downward displacement \u03b42 here",
            " The analysis could be \u201cextended\u201d in various ways for unilateral isplacement control (some of which have been reported in the litrature previously for force control), for loading at the crown or at ring joint. The possibility of member buckling could be included e.g., Fan et al., 2012; Hartono, 1997; Kani and McConnel, 1987; Ma t al., 2014; Yan et al., 2016b; Yang et al., 1997 ). For example, in artono (1997) for Dome I, the six members connecting the crown o the ring are assumed to have a slenderness ratio of 120, and the rst snap occurs at a bifurcation point with a force 20% lower than hat at the first limit point B in Fig. 3 . Nonlinearly elastic and elastic-plastic behavior could be anayzed (e.g., Blandford, 1996a,b; Handruleva et al., 2012; Kato et al., 003; Ramesh and Krishnamoorthy, 1994; Thai and Kim, 2009; Y G e V ( e p e S e 2 T o f H e J e f ( W p c a P t t A t t t a R A A A B B B C C C F F F F G G H H H H H H H H J J K K K K K K K K K K L L L L L M M M M O P P P P u et al., 2011 ). The dynamics of snapping could be studied (e.g., omez et al., 2017; Pandey et al., 2014; Wiebe et al., 2013 ). The ffect of the velocity of pushing could be investigated ( Plaut and irgin, 2017 )",
            " The suport joints could be allowed to move in the support plane with an lastic resistance (e.g., Fujimoto et al., 1993; Kato et al., 2015 ). The effect of imperfections could be studied (e.g., Borri and pinelli, 1988; Chen and Shen, 1993; Fong et al., 2012; Kato t al., 1998; Kleiber and Wo \u0301zniak, 1990; Shon et al., 2013, 014; Su and Hai, 2016; Tanaka et al., 1985 ). For example, in anaka et al. (1985) for Dome I, the effect of raising the heights f ring joints 2 and 5 in Table 1 from 6.216 to 6.416 decreases the orce at the first limit point B in Fig. 3 by 8%. Finally, the joints could be flexible (semi-rigid) or fixed (e.g., an et al., 2016; Kani and Heidan, 2007; Kato et al., 1998; L\u00f3pez t al., 2007; Ma et al., 2016; Meek and Tan, 1984; Ramalingam and ayachandran, 2015; Xiong et al., 2017; Yamada et al., 2001; Yan t al., 2016a; Yang et al., 1997; Zhao et al., 2016 ). For comparison, similar loading of a two-member toggle rame ( Rezaiee-Pajand and Alamatian, 2011 ), a shallow arch Plaut, 2015b ), or a rigid-bar model of a shallow arch ( Croll and alker, 1972 ) would typically exhibit a single snap-through when ushed downward"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003508_0037549718784186-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003508_0037549718784186-Figure2-1.png",
        "caption": "Figure 2. Velocity diagram of MB. MB: mobile robot base.",
        "texts": [
            " The schematic diagram of the MM is shown in Figure 1, illustrating the different reference frames in the XY plane. The motion of the MB is considered with reference to the world coordinate frame {W}, the MB coordinate frame {G} is placed at the center of mass (CM) of the MB, and {E} is the frame of the tip of the MA. P is the point at which the MA is attached to the MB and {P} is the angular absolute frame parallel to {W}. The three-wheeled MB has been modeled by several researchers in the past.9\u201311 Figure 2 shows the velocities of the wheels and the CM of the MB. The kinematics are used here for bond graph modeling of the wheeled mobile robot. Vx, w, Vy, w and _f denote the linear velocities in the X and Y directions of the CM and rotational velocity about the vertical axis passing through the CM of the MB respectively in the world reference frame. _c1, _c2, and _c3 are the angular velocities of the wheels 1, 2, and 3 about the wheel axis respectively, and r is the radius of the three wheels. Mv represents the total mass (including the mass of the wheels) and Iv represent the moment of inertia of the MB (including the wheels) about the Z axis and I1, I2, and I3 are the moment of inertias of the wheels 1, 2, and 3 about the wheel axis, respectively. L is the radius of the MB. The kinematic equations relating the linear velocities of wheels with the CM velocities of the MB as shown in Figure 2 can be written as: _c1 _c2 _c3 2 4 3 5= 1 r sinf cosf L sinf1 cosf1 L sinf2 cosf2 L 2 4 3 5 Vx,w Vy,w _f 2 4 3 5, where, f1 = p 3 f and f2 = p 3 +f \u00f01\u00de The planar MA is appended to the MB at point P which is at a distance LP from the CM of the base. L1, L2, L3 and L4 are lengths of link-1, 2, 3 and 4 respectively, and u1, u2, u3 and u4 are corresponding joint angles, as shown in Figure 1. Since the MA is attached to the MB with an angular displacement f, the coordinates of each link are expressed in the frame {P} for development of the bond graph model of the MA"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002005_toh.2014.2330300-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002005_toh.2014.2330300-Figure3-1.png",
        "caption": "Fig. 3. Definition of variables for model identification. d is exaggerated for a clear view.",
        "texts": [
            " The recorded data are passed to a recursive least-squares algorithm for iterative Hunt-Crossley model parameter estimation [27]. We denote this model as f \u00bc HNT \u00f0x; _x\u00de. Note that we assume a nearly constant feedback behavior on the surface and thus only identify one model. The next task is to obtain the data for the inclusionembedded sample. In this case, the manual indentation is constrained by guiding forces to follow a line from the point on the surface pTs closest to the initial position of the spherical inclusion pT0 (see Fig. 3 for an overview). During indentation, we again record data magnitude triples as in the previous step. However, this time these are determined as projections onto the line from pTs to pT0. In addition to this we record the time-changing position of the inclusion pT and the position of the tool tip pH . The former is obtained via a positional tracking system. The system consists of two infra-red cameras (TrackIR; NaturalPoint, Inc.) and a small retroreflective marker and provides sub-millimeter-accuracy measurements of the center of the inclusion through triangulation",
            " Using these we identify another Hunt-Crossely model HT \u00f0l; _l\u00de, which describes the inclusion-only force response at the single contact point pTs, when probing into the direction of pT . The model will be used to determine feedback forces at arbitrary contact locations, as described in Section 5. The second step of the identification process is the characterization of inclusion movement in response to external forces. The task is to identify nonlinear changes of the inclusion displacement vector (d\u00f0t\u00de in Fig. 3) with respect to an external force due to indentation, fT \u00f0t\u00de, when there is only one contact. We again employ the Hunt-Crossley model for this relation. In general, the displacement of the inclusion depends on the boundary conditions (e.g., the rigid support beneath the sample) and varies by direction. As an approximation, we identify three independent Hunt-Crossley models for the three Cartesian directions. The process begins with determining d\u00f0t\u00de and fT \u00f0t\u00de during unconstrained palpation of the inclusion-embedded model with one tool"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002715_s0263574716000436-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002715_s0263574716000436-Figure1-1.png",
        "caption": "Fig. 1. Leader following formation control system.",
        "texts": [
            " Given a smooth bounded reference spatial trajectory [xd, yd, zd]T which is generated by an open-loop path-planner (the leader), torque signals [F1i , F2i , F3i]T for the ith follower UUV are designed under the following requirements: the ith follower UUV tracks the virtual leader such that limt\u2192\u221e |ei6 \u2212 ei6d | \u2264 \u03b5ei6, limt\u2192\u221e |ei7 \u2212 ei7d | \u2264 \u03b5ei7 and limt\u2192\u221e |ei8 \u2212 ei8d | \u2264 \u03b5ei8 where ei6, ei7 and ei8 are the range and bearing of the ith follower UUV with respect to the virtual leader (detailed in Fig.1) and ei6d , ei7d and ei8d denote their desired values, respectively. The terms \u03b5ei6, \u03b5ei7 and \u03b5ei8 are arbitrarily small positive constants. Remark 3. In this paper, we focus on the problems caused by the constraints on control inputs. We assume all follower UUVs have the identical kinematic model. The formation control problems in the presence of dynamics issues such as slipping and differences between the UUVs, which may be caused by sizes, masses and damage, are more complicated and will be investigated in our future research"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003911_s42417-019-00111-6-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003911_s42417-019-00111-6-Figure2-1.png",
        "caption": "Fig. 2 Bearing test rig of CWRU",
        "texts": [
            " (15)Outer-race defect: fOR = n 2 fr ( 1 \u2212 d D cos ) , (16)Inner-race defect: fIR = n 2 fr ( 1 + d D cos ) , Case 1: Detect the\u00a0Including Fault in\u00a0Rolling Bearing The bearing-testing signal from the bearing Data Centre in Case Western Reserve University (CRWU) is adopted in this case [21]. The fault vibration signals of rolling bearing were measured by an accelerometer that was equipped on the bearing support, which was connected to a torque transducer, and the torque transducer was connected to the dynamometer, the test bed was driven by the motor. The detail of test equipment is shown in Fig.\u00a02. The fault bearing is deep-groove ball bearing 6205- 2RSL JEM SKF, the dimensions and specifications are listed in Table\u00a01. The dimensions of including fault were 0.007 inch, the depth of all fault was 0.011 inch. And the four different load conditions with 0, 1, 2, and 3 horse power (hp) are measured, respectively. The sample frequency and duration of each signal were 12\u00a0kHz and 10\u00a0s, respectively. In this study, the signals are selected at speed of 1797\u00a0rpm (29.95\u00a0Hz) for 0 hp and 1730\u00a0rpm (28"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002465_amm.816.54-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002465_amm.816.54-Figure14-1.png",
        "caption": "Fig. 14. The values and spots of the concentration of contact pressures for the angular position 60 o",
        "texts": [
            " Figure 11 presents the state of the Mises stresses with the enlarged view of the place of the occurrence of maximum stresses. The values of the reaction forces and their components are also presented. The values and spots of the concentration of contact pressures are presented on Figure 12. The position 60 o . Figure 13 presents the state of the Mises stresses with the enlarged view of the place of the occurrence of maximum stresses. The values of the reaction forces and their components are also presented. The values and spots of the concentration of contact pressures are presented on Figure 14. The position 75 o . Figure 15 presents the state of the Mises stresses with the enlarged view of the place of the occurrence of maximum stresses. The values of the reaction forces and their components are also presented. The values and spots of the concentration of contact pressures are presented on Figure 16. Table 1 presents the list of component reactions and their resultant reactions. This table also shows the forces and moments of friction which are needed to disassemble the connection. Figure 17 illustartes the graphic representation of the Mises stresses and contact pressures in function of the angular position of the shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001956_citcon.2015.7122604-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001956_citcon.2015.7122604-Figure2-1.png",
        "caption": "Fig. 2: Unshielded, asymmetrical cabling also develops circulating currents that will run through the bearing.",
        "texts": [
            " This circulating flux generated in the stator gives a rotor flux alternating between symmetrical and asymmetrical. The flux changes induce a shaft voltage which has the same frequency as the supply frequency, or a multiple of it. The axial shaft voltage result in a low frequency, circulating current that flows through the bearings. The classic problem of flux asymmtries is compounded on larger 2 pole motors due to their size. These large motors inherently have greater flux asymmetries than smaller multi-pole motors. Figure 2. 978-1-4799-5580-0/15/$31.00 \u00a9 IEEE 2014-CIC-1045 A YSD motor operates using a frequency converter. A frequency converter is an excellent componnent of a motor. It simplifies design and control for variable speeds, reduces power consumption and enables high-speed designs. Frequency converters; however, have caused an increase in electrical damage of bearings. The reason for this is a power switching semiconductor used in the converter which is a gate turn off transistor (GTO) or more commonly an insulated gate bi-polar transistor (IGBT) which dominates the market today"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000683_icra.2013.6631258-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000683_icra.2013.6631258-Figure4-1.png",
        "caption": "Fig. 4. Reconfiguration during hand executing task r\u0308nd. JjM\u22121(JnM\u22121)+ \u00a8\u0303rnd is a induced acceleration of j-th link by \u00a8\u0303rnd. When \u22061r\u0308j is produced through 1l, 1r\u0308j appears at j-th link, meaning avoiding acceleration can be achieved as a secondary task.",
        "texts": [
            " (6), 1r\u0308j \u2212 J\u0307 j q\u0307 \u2212 J jM \u22121(JnM\u22121)+(r\u0308nd \u2212 J\u0307nq\u0307) =J jM \u22121[In \u2212 (JnM\u22121)+(JnM\u22121)] 1l (12) Here, we define three variables shown as 1 \u00a8\u0302rj = J\u0307 j q\u0307 + J jM \u22121(JnM\u22121)+(r\u0308nd \u2212 J\u0307nq\u0307) (13) \u2206 1r\u0308j = 1r\u0308j \u2212 1 \u00a8\u0302rj (14) 1\u039bj = J jM \u22121[In \u2212 (JnM\u22121)+(JnM\u22121)] (15) In Eq. (13), 1 \u00a8\u0302rj is represents acceleration caused by manipulator\u2019s shape change. In the right side of Eq. (13), the first term denotes Coliolis and centrifugal acceleration of jth link, the second term represents influence of 1 \u00a8\u0303rnd on j-th link except Coliolis and centrifugal acceleration at n-th link. Then, Eq. (12) can be rewritten as \u22061r\u0308j = 1\u039bj 1l (16) The relation between 1r\u0308j and \u22061r\u0308j is shown in Fig. 4. However, the problem is whether we can yield desired \u22061r\u0308jd, that is, whether we can find 1l to generate \u22061r\u0308jd. From Eq. (16), we can obtain 1l as 1l = 1\u039b+ j \u22061r\u0308jd + (In \u2212 1\u039b+ j 1\u039bj) 2l (17) In Eq. (17), 2l is an arbitrary vector satisfying 2l \u2208 Rn. Assuming that 1l is restricted as \u20161l\u2016 \u2264 1, then we obtain next relation, (\u22061r\u0308jd)T(1\u039b+ j )T1\u039b+ j \u22061r\u0308jd \u2264 1 (18) If rank(1\u039bj) = m, Eq. (18) represents an ellipsoid expanding in m-dimensional space, holding \u22061r\u0308jd = 1\u039bj 1\u039b+ j \u22061r\u0308jd, \u22061r\u0308jd \u2208 Rm (19) which indicates that \u22061r\u0308jd can be arbitrarily generated in m-dimensional space and Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure8-1.png",
        "caption": "Fig. 8. Cylindrical body resistant force on the surface changes under the water\u2013 vapor flow.",
        "texts": [],
        "surrounding_texts": [
            "Total energy equation for interphase heat transfer [5]: @ dt \u00f0raqaha;tot\u00de ra @p dt \u00fer\u00f0raqaUaha;tot\u00de \u00bc r\u00f0rakarTa\u00de \u00fe rar\u00f0Uasa\u00de \u00fe SEa \u00fe Qa; \u00f04\u00de where ha,tot, Ta, ka, qa \u2013 total enthalpy, temperature, thermal conductivity and density of phase a, ra \u2013 volume fraction of phase a, Uasa \u2013 viscous work term, Sea \u2013 external heat source, Qa \u2013 interphase heat transfer to phase a;C\u00feabhbs;tot C\u00febahas;tot \u2013 heat transfer induced by interface mass transfer. The homogenous heat transfer model was used in this research. Eq. (4) describes the expression of complete energy for nonhomogenous interphasic heat transfer; however, the expression of equation for homogenous model is identical. The only difference is that the coefficient of heat transfer is taken large enough to ensure an equal temperature of different phases."
        ]
    },
    {
        "image_filename": "designv11_13_0003134_dnx030-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003134_dnx030-Figure5-1.png",
        "caption": "Fig. 5. Plots of U(x) with {F\u0302, H, R\u0302, W} = {1, 1, 0.5, 0.2}. Plots (c) and (d) are enlarged views of (a) and (b), respectively. The yellow closed curve represents U(x) = U([\u03a3d , 0]T ), which bounds the subset D. The blue curves are subsets on which U(x) is indifferentiable. The green curves are the boundary of Y where U(x) = 0.",
        "texts": [
            "40) where U1(x) \u0394= G2 ([ W root \u03be\u2208[Vy ,+\u221e) (G1(x) \u2212 G1([W , \u03be ]T )) ]) = 2F\u03022H2 F\u0302 + R\u0302 \u239b\u239d(1 + \u221a G1(x) 2F\u03022H2/(F\u0302 \u2212 R\u0302) )2 \u2212 1 \u239e\u23a0 (3.41) U5(\u03c3p) \u0394= U1(x\u03c3 (\u03c3p)) (3.42) x\u03c3 (\u03c3p) \u0394= [ \u03c3p v\u03c3 (\u03c3p) ] \u0394= \u23a1\u23a3 \u03c3p \u03a3d \u2212 \u03c3p \u03a3d \u2212 W Vy \u23a4\u23a6 (3.43) and D1 \u0394= {x \u2208 D\\Y | sgn(v)\u03c3 \u2265 W \u2227 |v| \u2265 v\u03c3 (|\u03c3 |)} (3.44) D2 \u0394= { x \u2208 D\\Y | |v| \u2264 W \u2228 (sgn(v)\u03c3 \u2264 \u2212W \u2227 G2(sgn(v)x) \u2265 U5(|\u03c3 |))} (3.45) D3 \u0394= (D\\Y)\\Int(D1 \u222a D2). (3.46) Here, Di are chosen so that they overlap each other at their boundaries and that Di \u2229 Y = \u2205 and D\\Y = D1 \u222a D2 \u222a D3 are satisfied. (Figure 5 illustrates these sets and the function U(x).) The function U(x) is positive definite with respect to the set Y and is continuous in D. Therefore, the proof can be completed by showing that, for any initial states x(0) = x0 \u2208 D\\Y , there exists a positive constant q(x0) such that D\u2217 t U(x) \u2264 \u2212q(x0) for all t > 0. (See, e.g. Polyakov & Fridman (2014, Theorem 11).) If x \u2208 D1 \u2282 cl(R1 \u222a R3), in which v = 0, the following is satisfied: U\u0307(x) = U\u03071(sgn(v)x) = F\u0302 \u2212 R\u0302 F\u0302 + R\u0302 \u239b\u239d\u221a 2F\u03022H2 G1(sgn(v)x)(F\u0302 \u2212 R\u0302) + 1 \u239e\u23a0 G\u03071(sgn(v)x) \u2264 F\u0302 \u2212 R\u0302 F\u0302 + R\u0302 G\u03071(sgn(v)x) \u2264 \u2212\u03b1|v| < 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000413_s12206-011-1201-6-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000413_s12206-011-1201-6-Figure2-1.png",
        "caption": "Fig. 2. Configuration of bearing-rotor system of the air blower.",
        "texts": [
            " The rotor was also analyzed by FEM and results were verified, which shows that the new method proposed for vibration characteristics calculation of air blowers in this paper is credible. Fig. 1 shows an air blower used in a power station in China. It is mainly composed of a motor, pulley and bearing-rotor. The most important part of the blower is its rolling bearingrotor system. During the vibration analysis of bearing-rotor system, the foundation is treated as rigid and its vibration is ignored. Fig. 2 shows the geometry parameters of the rolling bearing-rotor system, which is composed of an impeller, pulley, shaft and a pair of tapered roller bearings (30218). The impeller, located in the middle of the two bearings, is composed of several bar pieces of weight 264.45 kg and inertia 37.49 2kg mi ; the pulley located on the right side of the shaft has a weight of 158.97 kg with inertia 3.27 2kg mi and the rotor\u2019s working speed is 6000 rpm. According to lumped mass method, the beam is simplified to be several disks and beams, and the disk is considered to have no thickness and the beam is considered to have no mass"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001136_s00170-017-0625-2-Figure16-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001136_s00170-017-0625-2-Figure16-1.png",
        "caption": "Fig. 16 Origin of incomplete fusion defect generated in ladder block samples",
        "texts": [
            " With the additive process of ladder block, the chaotic extent of grains near the surface decreases gradually, and the grains on the surface instead become more directional with the deposition process. The fine columnar grains that point to the scanning direction dominate the surface finally. The internal defects of the pulsed-laser deposition with metal wire include incomplete fusion, cracking, and porosity. Figure 15 displays the typical incomplete fusion defects, which are considered to be the most common defect in the depositions formed by a pulsed-laser input and wire feed, particularly for the ladder-block structures. Figure 16 illustrates the origin of incomplete fusion defects generated in ladder block samples (the arrows show direction of filling up of the liquid metal). The incomplete fusion defect is prone to be produced by an overlap process between the nearby beads or layers in the deposition process. Since the top surface of each layer of ladder blocks always fluctuates, the liquid metal has difficulty in filling up an uneven surface in a very short solidification time. Moreover, in the pulsed laser and wire-based deposition processes, the molten pool is only 2\u20133 times greater than the diameter of a fine wire"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002445_ijmmme.2016010103-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002445_ijmmme.2016010103-Figure2-1.png",
        "caption": "Figure 2. Gearbox assembly of 4-speed transmission gearbox",
        "texts": [
            " The designing and assembly were done using solid Edge (Solid Edge, 2006) and Pro-E (Pro-E, 2013). All designing parameters were obtained from measurements and drawing sheets. The casing encloses the gear assembly, input, output and lay shaft. In this study only direct drive 4-speed gears were considered. Figure 1 shows the full assembly of transmission gearbox of vehicle. The isometric view of vehicle shows casing designed in three parts. The main part covers the gear assembly and other two parts used to enclose the transmission casing. Figure 2 shows 4-speed gear assembly and shafts for deformation analysis. The numerical simulation was performed for the transient structural, thermal analysis of gear train under the influence of load, rotational speed and convection heat transfer coefficient. Gear oil works as gearbox lubricant to cool the gears and transfer the heat through convective process. The overheating of gears reduces the efficiency and life span of gears. The transmission gearbox assembly replacement is costly, so this study focus the thermal prospectus of gearbox so that the life span of a transmission gear train can be increased"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002640_j.measurement.2016.05.021-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002640_j.measurement.2016.05.021-Figure1-1.png",
        "caption": "Figure 1. Forming principle of TP worm",
        "texts": [
            " Compared with the cylindrical worm, it has the advantages of more meshing teeth and the double-line contacting worm flank. All these qualities bring it high load capacity and enhance its significant role in lightweight gear equipment. However, it is difficult to manufacture the worm with high quality to guarantee its excellent performance. The measurement of the complicated geometry is critical to solve the problem. Cylindrical worm measurement is relatively fully developed [3], while TP worm measuring The model of the worm is based on the forming principle of the worm, as shown in Figure 1. The plane in which the grinding wheel is placed is the generating plane. The grinding wheel rotates around the basic circle and the worm rotates around the axis of itself at the speeds of 2 and 1 , respectively. 1 2/  is constant. The whole process could be regarded as the contact line of the plane is copied to the flank of the worm after coordinate transformation [7]. Figure 2. Coordinate of the TP worm and the worm gear The coordinate systems of the worm and the worm gear are placed as shown in Figure 2 [8, 9]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001042_0959651817698350-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001042_0959651817698350-Figure12-1.png",
        "caption": "Figure 12. Manipulator configuration stage II.",
        "texts": [
            " The whole manipulator is divided into two sections for simplicity. In the first section of the manipulator, only joint 1 is active, and in the second section, only joint 4 is active. At this stage, d1LO \\ 0 for obstacle 1 is assumed in section 1 and d2LO \\ 0 for obstacle 2 is assumed in section 2 of the manipulator. These mean that the robot system suffers a serious collision from both the obstacles. Now, assume that the planar space robot system avoids both the obstacles one by one. Stage II. This stage is schematically represented in Figure 12. Figure 12 shows the circumstance (assuming) when the robot avoids obstacle 1 but unable to avoid obstacle 2. The occurrence of this situation is discussed as follows: suppose link 2 tip comes close to the obstacle 1 during the task execution. If it continues, the first section of the manipulator will collide with the obstacle 1. To avoid this collision, let joint 1 is made fixed during manipulation at minimum positive value of d1LO ensuring a safe margin between the barrier and the surrounding links. This act will restrict motion of its influencing link tip, that is, tip of link 2 as joints 2 and 3 are already at FJPs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003283_0954406217745336-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003283_0954406217745336-Figure2-1.png",
        "caption": "Figure 2. Force diagram of bearing\u2013wheel\u2013rail system: (a) Force diagram of left bearing; (b) Force diagram of right bearing; (c) Force diagram of wheel; (d) Force diagram of rail.",
        "texts": [
            " The bearing\u2013wheel\u2013rail system consisted mainly of an adapter, bearing, wheel, rail axle, and rail, as shown in Figure 1. The adapter is directly fitted to the bearing, the bearing inner ring and axle are interference fitted, and the wheels are directly fitted onto the corresponding position of the axle, so no slippage was allowed during the operation. The wheel is in direct contact with the rail. The force distribution of the system was calculated according to its structure and working principle and is shown in Figure 2. The adapter was installed between the railway bearings and bogie. It bore the bogie\u2019s impact load and passed that impact load on to the bearing. Most railway bearings used double rows of circular cone roller bearings, which could withstand vertical and horizontal loads. This paper analyzed the bearing\u2019s strong nonlinear coupling characteristics, taking the bearing outer ring as a node, and established its own vibration equation through the analysis of the mechanical characteristics. The force distribution of the left and right adapter and bearings is shown in Figure 2(a) and (b). mc is the mass of the adapter system; kcy is radial stiffness of the adapter; ccy is the radial damping of the adapter; Fv1(t), Fv2(t) are the right and left adapter\u2019s vertical load time histories; dc1, dc2 are the right and left adapter\u2019s radial displacements; db1, db2 are the right and left bearing outer rings\u2019 radial displacements; mb is the mass of the bearing outer ring; Fr1, Fr2 are the radial nonlinear force of the right and left bearings. Vibration equation. The dynamics equation was determined using the stress of the right adapter\u2013bearing mc \u20acdc1 \u00fe kcy\u00f0dc1 db1\u00de \u00fe ccy\u00f0 _dc1 _db1\u00de \u00bc FV1 \u00f0t\u00de \u00femcg \u00f01\u00de The right bearing outer ring\u2019s dynamics equation is mb \u20acdb1 kcy\u00f0dc1 db1\u00de ccy\u00f0 _dc1 _db1\u00de \u00bc mbg Fr1 \u00f02\u00de The dynamics equation can be acquired according to the stress of left adapter\u2013bearing mc \u20acdc2 \u00fe kcy\u00f0dc2 db2\u00de \u00fe ccy\u00f0 _dc2 _db2\u00de \u00bc FV2\u00f0t\u00de \u00femcg \u00f03\u00de The left bearing outer ring\u2019s dynamics equation is as follows mb \u20acdb2 kcy\u00f0dc2 db2\u00de ccy\u00f0 _dc2 _db2\u00de \u00bc mbg Fr2 \u00f04\u00de During the operation, the adapter bore the vertical force, which it passed onto the bearing outer ring; hence, the bearing outer ring also bore the vertical load, horizontal load, and torque load",
            " The rail wheel\u2019s contact stress is calculated using equation (25) \u00f0t\u00de \u00bc S N\u00bd 1=3 \u00f025\u00de S \u00bc 2:49R 0:251 107 TB\u00f0 \u00de 1:49R 0:376 107 LM\u00f0 \u00de \u00f026\u00de where S(N2/3/m2) is the stress contact defined using the classic Hertz theory. The range of R was 0.15 0.6m. TB was the conical tread wheel and LM was the wear type wheel tread. The long pillow embedded type rail system, which has no frantic jumble, served as an example. The system was mainly composed of steel, fasteners (rail pads), concrete pillows, concrete road bed boards, and bases. The rail\u2019s elastic deformation was mainly provided by rail pads. The system\u2019s mechanics model is shown in Figure 2, where kp is the stiffness of the rail pads; cp is the damping of the rail pad, which is acquired by experiments. Rail systems are very complicated. Actual numerical computation always involves a simplified supported beam model with finite length. Practical results have shown that such processes can completely satisfy the precision demand of engineering. This paper uses the Euler beam,24\u201326 and the force model is shown in Figure 6, where P is the rail\u2019s contact force, the right wheel\u2019s is N0Ry. The wheels\u2019 speed is v; FRi(i\u00bc 1,"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure1-1.png",
        "caption": "Fig. 1. (a) Real bearing (b) DEM bearing and its geometrical characteristics.",
        "texts": [
            " To distinguish this new description from multibody and FEM approaches, a discrete element approach is proposed for a bearing application [4,5] . Usually ignored or difficult to model, some important mechanical considerations are treated in this paper. The cage is made of discrete elements moving along the pitch diameter ( R inner + R outer 2 ) and interacting with rolling elements. The flexibility of rings is taken into account and internal clearance can be controlled. A comparison between a real bearing and its numerical discrete model is proposed in Fig. 1 . In 2D simulations, the gyroscopic forces are not taken into account. By using the smooth Discrete Element Method (DEM), initially developed by Cundall and Strack [27,28] , the contact forces in a bearing are described with a model depending on an elastic force displacement law, Coulomb\u2019s friction and viscous damping. The contact occurs only when discrete elements with radii a i and a j interpenetrate, which means that a contact between a rolling element and a ring or a contact between a rolling element and a cage component exists as suggested by the Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003448_j.procir.2018.03.277-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003448_j.procir.2018.03.277-Figure4-1.png",
        "caption": "Fig. 4 Photograph of vise no. 2. Made by design students.",
        "texts": [
            " Findings There were three groups that chose the optional assignment with 3D-printing. One of the groups were design students, and two of the groups were mechanical engineer students. The students chose the group members themselves. 3.1. Practical results of the process The assignment was not carried out exactly as planned. The students were supposed to get the nylon parts before they started, but because the nylon printer had a failure and needed to be repaired, the students got the parts at the end of the project period. These are the white parts in Fig. 4 and Fig. 5. Nevertheless, all the groups managed to complete the assignment within the deadline, and ended up with a fully functional vise. The vices made by the mechanical engineer students are shown in Fig. 3 and Fig. 5. The vise made by the design students is shown in Fig. 4. The design students used this unforeseen opportunity to experiment and then print the missing parts in plastic with the cheap desktop printers, as shown in Fig. 6. They also managed to print the small screws in plastic, although steel screws were handed out (see the two screws at the bottom in Fig. 6). Beforehand it was considered unlikely that the threaded parts could be printed in plastic, but the design students did it anyway. They did however experience that the smallest plastic screws broke during assembly because they were not strong enough, and the threads on the spindle (see third part from the top in Fig",
            " Findings There were three groups that chose the optional assignment with 3D-printing. One of the groups were design students, and two of the groups were mechanical engineer students. The students chose the group members themselves. 3.1. Practical results of the process The assignment was not carried out exactly as planned. The students were supposed to get the nylon parts before they started, but because the nylon printer had a failure and needed to be repaired, the students got the parts at the end of the project period. These are the white parts in Fig. 4 and Fig. 5. Nevertheless, all the groups managed to complete the assignment within the deadline, and ended up with a fully functional vise. The vices made by the mechanical engineer students are shown in Fig. 3 and Fig. 5. The vise made by the design students is shown in Fig. 4. The design students used this unforeseen opportunity to experiment and then print the missing parts in plastic with the cheap desktop printers, as shown in Fig. 6. They also managed to print the small screws in plastic, although steel screws were handed out (see the two screws at the bottom in Fig. 6). Beforehand it was considered unlikely that the threaded parts could be printed in plastic, but the design students did it anyway. They did however experience that the smallest plastic screws broke during assembly because they were not strong enough, and the threads on the spindle (see third part from the top in Fig",
            " An example of one of the drawings is shown in Fig. 7. It is not possible to know how the quality of the drawings would have been if the students had not taken the optional additional work with 3D-printing. However, it should be noted that the 14 students participating had from none (0!) up to max three comments on their set of 2Ddrawings, which is an indication of an increased level of understanding when comparing with those students not taking on this task. Fig. 3 Photograph of vise no. 1. Made by mechanical engineer students. Fig. 4 Photograph of vise no. 2. Made by design students. Fig. 5 Photograph of vise no. 3. Made by mechanical engineer students. Fig. 6 The additional parts printed by the design students. Author name / Procedia CIRP 00 (2017) 000\u2013000 5 3.2. Cognitive result; students\u2019 learning outcome Results are presented in accordance with the structure of the interview guide presented in 2.2. The project as a whole: The two engineering groups both said that the time allocation was more than sufficient, and had it not been for printing time, this part of the assignment could have been completed in one day"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001988_3dp.2014.0017-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001988_3dp.2014.0017-Figure2-1.png",
        "caption": "Figure 2. SIS - metal machine ( left ) and block diagram of the machine ( right ).",
        "texts": [
            " The speed of SIS would be significantly higher if vector (instead of raster) printing is used. \u2022 SIS - metal parts are uncontaminated by any sort of binder or residue. The sintering characteristics of pure metal powder are superior to those of contaminated powder. \u2022 Being free of binder, sintering of SIS parts does not contaminate the sintering furnace, if mineral inhibitors such as metal salts are used. \u2022 There is no extra effort needed in support removal since any needed supports are attached to the \u201c sacrificial mold \u201d and not the part itself. The SIS - metal inkjet machine ( Figure 2 ) was developed at the Center for Rapid Automated Fabrication Technologies (CRAFT) 16 at the University of Southern California. This printer utilizes micropiezo technology that ejects droplets by mechanical movement of piezo actuators. A piezo printhead was chosen for the SIS - metal process because it provides more flexibility in the types of fluids that can be printed. This allows for 154 3D PRINTING MARY ANN LIEBERT, INC. \u2022 VOL. 1 NO. 3 \u2022 2014 \u2022 DOI: 10.1089/3dp.2014.0017 a wider range of candidate inhibitor solutions to be tested. The SIS - metal machine, developed in this research, consists of a three - axis motion system, a piezo - electric printhead, a controller board, motion control software, and user interface software (Figure 2). Parts were first designed and saved into STL format in computer aided design (CAD) software. Open - source and free software were then used to view, manipulate, and heal the STL files when necessary. The layered images were constructed using a predetermined uniform layer thickness of 120 \u03bc m for a given binary STL file format input. A dedicated graphical user interface was developed in Visual C# for STL slicing and image generation in the SIS process. SIS sliced image generation differs significantly from other AM processes because only the periphery of the part is marked for inhibitor deposition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000709_0278364913498909-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000709_0278364913498909-Figure1-1.png",
        "caption": "Fig. 1. Mixed sensor network model.",
        "texts": [
            " Extensive simulation results are presented in Section 5 which demonstrate several tradeoffs involved in the design parameters of the proposed algorithms and evaluate the performance of the proposed framework. Section 6 briefly discusses the relevant literature, and finally the paper concludes with Section 7 where conclusions are drawn and future plans are described. In this section we present the main modeling assumptions made throughout the paper and define some key performance indices such as area coverage and event detection. We consider a mixed sensor network made of a large number of sensor nodes deployed in a large region A as shown in Figure 1. We assume that the region to be monitored is a large rectangular area A and a large set S with S = |S| static sensor nodes that are deployed (typically randomly) in A, at positions xi = ( xi, yi), i = 1, . . . , S. In addition, we assume that a small set M of M = |M| mobile sensor nodes are available and their position after the kth time step is xi( k) = ( xi( k) , yi( k) ), i = 1, . . . , M , k = 0, 1, . . .. For notational convenience, we define the set of all sensor nodes N = S\u222aM and in this set the mobile nodes are re-indexed as m = S + 1, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003778_1.4042636-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003778_1.4042636-Figure1-1.png",
        "caption": "Fig. 1 Outside view of 4-UPS-UPU parallel mechanism",
        "texts": [],
        "surrounding_texts": [
            "2.1 Eccentricity Between Joint Elements With Clearance. As shown in Fig. 2, the fixed coordinate system OA XAYAZA denoted as fAg is set up on the fixed platform of 4-UPS-UPU parallel mechanism. The moving coordinate system OB XBYBZB denoted as fBg is set up on the moving platform. The branch chain coordinate system Ci XCiYCiZCi denoted by fCig is set up on the driving limbs, where i\u00bc 1, 2, 3, 4, and 5. When the clearances are not considered, the center of the sphere is coincident with the center of the ball sleeve and the degree-of-freedom is 3. It is assumed that there is a clearance at the spherical joint that between the third driving limb and the moving platform. The center of the sphere is not coincident with the center of the ball sleeve due to the existence of clearance, which causes the sphere can freely move in the ball sleeve and the spherical joint has six degrees-of-freedom. Constraint of three contact forces replaced constraint of three original motions. As shown in Fig. 3, S3 is the center of the ball sleeve, b3 is the center of the sphere, RS3 is the radius of the ball sleeve, Rb3 is the radius of the sphere, g is the contact plane of the two component of the spherical joint with clearance, and Ann and Att represent normal vector and tangent vector, respectively. In this study, \u201c*\u201d is used to distinguish the parameters with clearance. The coordinate of the center point of moving platform in fixed coordinate system fAg in ideal status can be expressed as APBO \u00bc \u00bdAPBOx APBOy APBOz T (1) Ideally, the rotation matrix that the fBg relative to fAg is written as A BR 5 cb 0 sb sasb ca sacb casb sa cacb 2 4 3 5 (2) 041010-2 / Vol. 14, APRIL 2019 Transactions of the ASME Downloaded From: https://computationalnonlinear.asmedigitalcollection.asme.org on 02/16/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use The coordinate of the center point of the moving platform when considering the spherical joint clearance in fAg can be expressed as AP BO \u00bc \u00bdAP BOx AP BOy AP BOz T (3) Rotation matrix that fB g relative to fAg can be expressed as A BR \u00bc cb 0 sb sa sb ca sa cb ca sb sa ca cb 2 4 3 5 (4) Eccentricity vector of the clearance is obtained as Ae\u00bc APb3 AP S3\u00bc APBO \u00fe A B RBPb3 AP BO A B R BPS3 (5) Eccentricity scalar of the clearance is expressed as e \u00bc ffiffiffiffiffiffiffiffiffiffiffi AeTAe p \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Ae\u00f01; 1\u00de2\u00fe Ae\u00f02; 1\u00de2\u00fe Ae\u00f03; 1\u00de2 q (6) Normal unit vector of the contact plane g is written as Ann \u00bc Ae e (7) 2.2 Relative Normal Velocity and Relative Tangential Velocity Between Joint Elements. Whether the sphere and the ball sleeve are in contact or not can be expressed as d \u00bc e c (8) d < 0; no contact d \u00bc 0; start contact or start separation d > 0; contact and elastic deformation 8>< >: Where c \u00bc RS3 Rb3 is clearance value. As shown in Fig. 3, C3 on the ball sleeve and D3 on the sphere are contact points when the sphere and the ball sleeve are in contact, so the position coordinate of this two points in {A} can be written as APc3\u00bc AP S3 \u00fe Rb3 Ann (9) APD3\u00bc APb3 \u00fe Rb3 Ann (10) Then, the velocity of contact points can be expressed as AVC3\u00bc AV S3 \u00fe RS3 A _nn (11) AVD3\u00bc AVb3 \u00fe Rb3 A _nn (12) Where AV S3\u00bc Ax B Ar S3\u00fe AV BO (13) AVb3\u00bc AVS3\u00bc AxB ArS3\u00fe AVBO (14) The relative normal velocity AVn and the relative tangential velocity AVt of joint elements with clearance can be obtained by projecting the relative velocity of contact points onto contact plane and normal plane, respectively. So we can get AVn \u00bc \u00bd\u00f0AVD3 AVC3\u00deTAnn Ann (15) AVt \u00bc \u00f0AVD3 AVC3\u00de AVn (16) The scalar of relative tangential velocity is expressed as Vt \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi AVT t AVt q (17) Owing to AVt \u00bc VA t tt, the tangential unit vector of the contact plane g is written as Att \u00bc AVt Vt (18)"
        ]
    },
    {
        "image_filename": "designv11_13_0001101_j.procs.2017.05.324-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001101_j.procs.2017.05.324-Figure2-1.png",
        "caption": "Fig. 2. Finite element model of (a) sun-planet gear pair; (b) ring-planet gear pair.",
        "texts": [
            " In this section, a prototype of planetary gear train in a wind turbine drivetrain is taken as an example to demonstrate the effects of tooth crack on fault signals. The basic design parameters of the example system are listed in Table 1. The pressure angles of sun, planet and ring gears are 20 deg. The mesh damping of sun-planet pair and ring-planet are set as 242.6 Ns/m and 410.3 Ns/m, respectively. To formulate the time-varying mesh stiffness accurately, a three-dimensional finite element model for the sunplanet and the ring-planet gear pairs is developed in ABAQUS/Standard. As shown in Fig. 2, the whole body of the gears in mesh are modelled with linear brick element C3D8R. The tooth conjunction is modelled as general contact which includes elastic Coulomb frictional effects. Two reference points are created at the centers of input gear and output gear, which are connected to the gear hubs with coupling constraints. The output gear is fixed and the input gear is set to rotate only about its rotational axis. A torque is applied at the rotational axis of the input gear to make the input and output gear tooth surfaces contact with each other"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000378_j.protcy.2013.04.050-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000378_j.protcy.2013.04.050-Figure1-1.png",
        "caption": "Fig. 1. Simple diagram of the gas metal arc welding process (Naidu, 2003)",
        "texts": [
            " The objective of the first part of this study is to find the optimal bead area geometry in the GMAW process. The first part of this paper gives a brief introduction to the GMAW process, and then a statistical threelevel factorial analysis is described in section 3. Finally a radial basis function neural network is used for the prediction of the cross sectional area of the welding bead and some conclusions are presented. Welding is a multi-energy process that involves various physical and chemical phenomena, such as plasma physics, heat flow and fluid and metal transfer and heat. GMAW welding (Fig. 1) establishes an electrical arc between a continuously fed electrode and the weld pool; this is protected by a gas administered externally. The heat generated by the arc melts the base metal surface as well as the electrode tip. The molten electrode is transferred to the workpiece through the electric arc and serves as the filler metal (weld bead). 2.1 Development of experimental design The chosen factors for the design of experiments were the wire feed speed (an indirect measure of the arc current), the welding voltage and the welding travel speed, as output we observe the cross sectional area of the welding bead"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003467_s40997-018-0184-7-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003467_s40997-018-0184-7-Figure4-1.png",
        "caption": "Fig. 4 The coordinate system of TCA for a helical gear pair",
        "texts": [
            " By coordinate transformation, the tooth surface equation of pinion/gear in coordinate system oixiyizi is described as ri \u00bc xi yi zi 1 2 664 3 775 \u00bc Mcoirc \u00f04\u00de where Mcoi \u00bc cos/i sin/i 0 0 sin/i cos/i 0 0 0 0 1 0 0 0 0 1 2 664 3 775 1 0 0 ri 0 1 0 si 0 0 1 0 0 0 0 1 2 664 3 775; /i \u00bc si=ri According to the Willis law, the rack cutter tooth surface envelops the gear tooth surface in the process of machining; this process must satisfy the following meshing equation: f \u00f0ui; li;/i\u00de \u00bc n~ci v~\u00f0ci;i\u00de ci \u00bc 0 \u00f05\u00de By solving formula (5), li can be replaced by ui and /i, i.e., li \u00bc li\u00f0ui;/i\u00de \u00f06\u00de The coordinate system of TCA for a helical gear pair is shown in Fig. 4. By coordinate transformation, the tooth surface equation of pinion in coordinate system ofxfyf is described as r1f \u00bc x1f y1f z1f 1 2 664 3 775 \u00bc Mo1oh1Moh1oh2Moh2of r1 \u00f07\u00de where Mo1oh1 \u00bc cos/1 sin/1 0 0 sin/1 cos/1 0 0 0 0 1 0 0 0 0 1 2 6664 3 7775 Moh1oh2 \u00bc 1 0 0 E \u00fe DE 0 1 0 0 0 0 1 0 0 0 0 1 2 6664 3 7775 Moh2of \u00bc 1 0 0 0 0 cos dr sin dr 0 0 sin dr cos dr 0 0 0 0 1 2 6664 3 7775 By coordinate transformation, the tooth surface equation of gear in coordinate system ofxfyf is described as r2f \u00bc x2f y2f z2f 1 2 664 3 775 \u00bc Mo2of r2 \u00f08\u00de where Mo2of \u00bc cos/2 sin/2 0 0 sin/2 cos/2 0 0 0 0 1 0 0 0 0 1 2 664 3 775 In coordinate system ofxfyf, the contact point of two tooth surface must have a common position vector and a common normal vector, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003514_s11661-018-4741-x-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003514_s11661-018-4741-x-Figure1-1.png",
        "caption": "Fig. 1\u2014(a) The CAD model cross-section, (b) as-printed views of the metal hybrid structure, and (c) laminography setup showing the incident neutron beam, the sample on a rotation stage and the detector. (a) The light gray circles indicate the Co-Cr alloy and the dark gray areas are the IN718 alloy. The boundaries of these features are apparent in the as-printed sample (b). Also, shown in (a) are the locations where the EDS (red rectangle) and XRD (1, 2, 3, 4) measurements were performed (Color figure online).",
        "texts": [
            " Due to the non-availability of elemental powders, a Co-Cr alloy and Inconel-718 (IN718) were used to fabricate a composite material. We realize that this route introduces Cr into the Fe-Ni-Co system; however, the Cr in both alloys are in a solid solution and is expected to have a minimal effect, if any. The compositions of the alloys used as well as the base plate are presented in Table I (all compositions are in wt pct). The metal hybrid model structure was designed around an IN718 matrix (dark gray, Figure 1(a)) housing 5 interlinked circles of the Co-Cr alloy (light gray, Figure 1(a)). The interlinked circles provide geometrical complexity and create more internal surface area. The sample was designed in the form of a thin hockey puck with a 49-mm diameter and ~ 2 to 3-mm thickness to facilitate non-destructive characterization, post-manufacture. The sample was built using a DMD-103D, a commercially available co-axial blown-powder laser (a diode laser operating at 910 nm) deposition technique, at Oak Ridge National Laboratory\u2019s (ORNL) Manufacturing Demonstration Facility (MDF)",
            " METALLURGICAL AND MATERIALS TRANSACTIONS A The nCT measurements were performed at the CG-1D neutron imaging beamline,[20,21] located at the High Flux Isotope Reactor, ORNL. Here, a CCD camera in a light-tight box is attached to a photo-camera lens. The lens looks indirectly at a LiF/ZnS scintillator, where the radiograph is formed, through a mirror positioned at a 45 deg angle with the neutron beam. Given the strong neutron attenuation of the specimen, a laminography approach was chosen for the scans where the sample was tilted 20 deg with respect to the neutron beam (see Figure 1(c)). Flat specimens with laterally extended dimensions can be challenging to measure using conventional computed tomography (CT) because projection data cannot be acquired from angles where the object is too thick (high absorption). Neutron laminography is a neutron-computed tomography where the sample\u2019s rotation axis is inclined with respect to and toward the beam normal. Drawbacks of the technique are the variation of voxel size throughout the sample and coarser spatial resolution than conventional CT. For these measurements, the L/D ratio was approximately 400 (where L is the distance between the beamline pinhole, of diameter D, and the detector). The nCT scan was measured between 0 and 360 deg (angle x in Figure 1(c)), with a 0.25 deg step-size and a 32 seconds exposure time for each radiograph. Given the complex geometry of the laminography scan, simple reconstruction methods such as the filtered back-projection developed in the context of parallel beam systems were not directly applicable. While analytic algorithms exist,[22] they can result in significant artifacts due to the noise in the system and the under-sampling in Fourier space due to the scanning geometry.[23] Hence, we have developed a model-based iterative reconstruction (MBIR) approach that models the complex geometry and casts the reconstruction as the solution to a regularized inverse problem",
            " The post-processing of the collected data was performed using the TSL OIM Analysis software and no additional data clean-ups were performed. Material Fe Ni Cr Co Al Nb C Mn Ti Mo Si Substrate rest nil nil nil nil nil 0.08 1.0 nil nil 0.8 IN718 rest 52 20 1 0.6 4.7 0.07 0.35 0.5 nil 0.3 Co-Cr nil nil 25 67 nil nil nil nil nil 8.0 nil *JEOL USA, Inc. 11 Dearborn Road, Peabody, MA 01960. METALLURGICAL AND MATERIALS TRANSACTIONS A The as-fabricated sample still attached to the build plate is shown in Figure 1(b). The sample was cut from the base plate using electric discharge machining and metallurgically polished prior to characterization. EDS elemental maps of Co, Ni, Fe, and Cr obtained from between the outer border of a Co-Cr cylinder and the IN718 matrix (as represented with the red rectangle in Figure 1(a)) are presented in Figure 2. Here the four elements are colored differently in their respective maps (blue for Co, yellow for Ni, green for Fe, and red for Cr) and higher color intensities correspond to increased elemental concentration in a given point and vice versa. The maps were not taken at equal distances from each other; nevertheless, they clearly outline the two zones with distinct Co and Ni, Fe-rich areas. Since Cr is present in both alloys (see Table I), its elemental map is relatively uniform, although a slight distinction can still be observed (see that the Cr intensities are slightly higher inside the Co-Cr ring)",
            " Figure 2 also reveals that the boundaries are not sharp with some inter-diffusion of alloying elements observed at the interfaces. This is expected to occur due to re-melting of the interfaces during the deposition of subsequent lines and/or layers. Based on this, it should be borne in mind that process parameters and materials should be selected appropriately to avoid any unexpected intermediate-phase formation, such as intermetallic compounds, at the interfaces. XRD experiments were performed at four different locations (indicated with numbers on Figure 1(a)) on the metal hybrid sample. Both IN718 and the Co-Cr alloys have FCC crystal structures, and thus, distinctions are expected to be minimal, if any. Also, due to the X-ray beam being 10 mm long, it is expected that in METALLURGICAL AND MATERIALS TRANSACTIONS A several of these scans, peaks from both Co-Cr and IN718 will be registered, i.e., interface regions will be included. In Figure 3, all the patterns show a single FCC phase with very similar lattice parameters (ranging from a = 3.613 to 3",
            " Scale bars are not presented in the 3D images as dimensions can be highly viewing-angle dependent. The reader is referred to the 2D slices for scales. Finally, Figure 4(d) shows 2D attenuation-based slices starting from the bottom and going to the top section of the sample. While the as-printed geometry agrees with the CAD model, the interfaces of the internal structure show some roughness. The magnified view, presented in Figure 4(c), shows the non-uniformities along the interfaces in more detail. This agrees well with the macroscopic melt lines seen in Figure 1(b) and the inter-diffusion observed in the EDS maps, as seen in Figure 2. However, this is not necessarily detrimental, except for cases where strict tolerances may be required, because the rough edges can also provide a more interlocked boundary. Overall, the Co-Cr internal structure is found to be continuous through the thickness of the hybrid component without structural discontinuities such as cracks or dissolved boundaries. The 2D slices, presented in Figure 4(d), show the continuity of the metal hybrid structure through the build layers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003055_bio.3310-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003055_bio.3310-Figure1-1.png",
        "caption": "FIGURE 1 Schematic diagram of the FI\u2013CL system",
        "texts": [
            " Under stirring, 20 ml of the red\u2013brown solution was adjusted to pH 7 with HCl solution, and 100 ml of ethanol was added dropwise to the mixture. The solution was then added to a 10% MgSO4 solution, stirred for 20 min and kept at room temperature. The solution was centrifuged for 10 min at 6000 r.p.m. (rcf = 2390 g) prior to use and the supernatant was diluted with water for the subsequent experiment. CL kinetic curves were generated in a glass cuvette using the batch method and CL detection was conducted on a BPCL luminescence analyzer. A schematic diagram of the FI\u2013CL system used is shown in Figure 1. Solutions of CNDs, carrier (water), iodide and KMnO4 were inserted into flow lines and driven by peristaltic pumps. The CND solution (60 \u03bcl) was injected into the iodide standard or sample solution, and mixed with KMnO4. The whole mixture was then delivered into the CL cell. The CL emission was detected and amplified using a photomultiplier tube at a high voltage of \u2212600 V. The concentration of iodide could be quantified from the change in CL intensity. Seaweed samples were purchased from a local market, air\u2010dried at room temperature and grated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002679_0309324715614194-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002679_0309324715614194-Figure4-1.png",
        "caption": "Figure 4. Areas of contact and adhesive interaction with increasing load for (a, b) a separate asperity and (c, d) the wavy surface.",
        "texts": [
            " For the case of wavy surface, the maximum contact pressure is smaller, while areas of contact and adhesive interaction are wider than for the separate asperity. The calculations show that for small loads, for which only tips of the wavy surface are in contact with the viscoelastic foundation, the contact pressure distributions are similar to those for the separate asperity. As the load per asperity increases and the interaction becomes closer to the saturated contact, the difference of the results for two models significantly increases. Figure 4 shows the areas of contact (dark-grey regions) and adhesive interaction (light-grey regions) for the separate asperity (Figure 4(a) and (b)) and for the wavy surface (Figure 4(c) and (d)). The graphs (Figure 4(a) and (c)) correspond to the dimensionless load P/(Rg)=200 for which we have separate contact spots for both cases, only slightly different from each other. The graphs (b) and (d) correspond to the load P/(Rg)=550, at which contact spots begin to coalescence for the wavy surface (Figure 4(d)) and the contact and adhesion areas are significantly different in size and shape from those for one asperity (Figure 4(b)). The results of analysis show that for small loads, while contact spots are far away from each other, the pressure distribution and the shape of the contact regions for the wavy surface are similar to those calculated for one asperity sliding over the viscoelastic foundation. As the load increases, the contact regions for the wavy surface become closer to each other and finally coalescence into one region; in this case, it is important to take into account not only shape and mutual location of the asperity tips but also of the valleys between asperities, that is, use the model for a wavy surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure5-1.png",
        "caption": "Fig. 5. (a) Two-DOF and (b) three-DOF mechanisms used for illustrating elusive barriers. Links OP and PQ have unitary length: OP = PQ = 1 .",
        "texts": [
            " Then, each of these N p parts can be processed by an individual processor dedicated to search barriers among its nodes, following Algorithm 5 . Obviously, when decreasing steps \u03b8 j (used for sampling self-motion manifolds, line 5 of Algorithm 2 ) and t i (used for discretizing the task space, line 3 of Algorithm 5 ), the proposed method will estimate the workspace barriers more accurately (at the expense of higher computation times). However, the proposed method may be unable to detect some elusive barriers independently of how small \u03b8 j and t i are. Next, we will analyze the mechanisms shown in Fig. 5 to illustrate two types of elusive barriers that the proposed method may fail to detect. The first type of elusive barriers that we will analyze are generated by self-motion manifolds that only exist when the task coordinates are placed exactly on these barriers, such that these manifolds vanish when perturbing the task coordinates away from these barriers. In order to illustrate this situation, consider the 2-DOF mechanism of Fig. 5 (a), which can be considered as a redundant robot with joint coordinates ( \u03b81 , \u03b82 ) (position of joint Q ) and task variable t 1 (orientation of link OP ). For a given t 1 , the self-motion manifold of this robot in plane ( \u03b81 , \u03b82 ) is a circle C centered at P = ( cos t 1 , sin t 1 ) and with radius 1. In the absence of additional constraints, C has constant topology \u2200 t 1 . Thus, the workspace of task variable t 1 is the unit circle, without interior barriers or singularities. Assume now that, in order to avoid collisions, point Q cannot be in the horizontal region H of plane ( \u03b81 , \u03b82 ) defined by: 1 < \u03b82 < 2 (strict inequalities)",
            "99}, then barrier B is split into two unidirectional barriers B 1 and B 2 that enclose the arc B 1 B 2 of the unit circle for which sin t 1 > 0.99 (see Fig. 6 (e)). For t 1 inside this arc, there exist two disjoint self-motion manifolds, which are the parts of circle C outside H . Elusive barrier B is the limit of arc B 1 B 2 degenerating into a point. The perturbed barriers B 1 and B 2 would be detectable by the proposed method using a sufficiently small step t 1 , since the arc enclosed by them has non-zero length and will contain nodes of grid G. In order to illustrate another type of elusive barrier, consider the mechanism of Fig. 5 (b), which is obtained from the mechanism of Fig. 5 (a) by adding another link of length d between joint Q and the vertical prismatic leg of length | \u03b82 |. This new mechanism has three DOF, and it can be considered as a redundant robot with joint coordinates ( \u03b81 , \u03b82 , \u03b83 ) and task variables ( t 1 , t 2 ) (position coordinates of joint Q relative to fixed joint O ). The orientation angle \u03c6 of link d is a passive variable. For a given task point Q = (t 1 , t 2 ) = (0 , 0) = O, the self-motion manifolds in the 3D joint space ( \u03b81 , \u03b82 , \u03b83 ) are two circles obtained as the intersection between cylinder { (\u03b81 \u2212 t 1 ) 2 + (\u03b82 \u2212 t 2 ) 2 = d 2 } and two planes { \u03b83 = K 1 } and { \u03b83 = K 2 } , where K 1 and K 2 are the two solutions of \u03b83 in [ \u2212\u03c0, \u03c0 ] that satisfy the following equation: (t 1 \u2212 cos \u03b83 ) 2 + (t 2 \u2212 sin \u03b83 ) 2 = 1 (8) Fixed joint O = (0 , 0) in this example is an interior barrier of the workspace, as justified next"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002554_1.4033128-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002554_1.4033128-Figure12-1.png",
        "caption": "Fig. 12 Two arrangements of dimples: (a) enclosed dimple and (b) opened dimple",
        "texts": [
            " In the balancing case, when the dimple is created around the center of the lubricated area, no solution of the equilibrium of moment is obtained, as shown in the previous papers [2,3]. For the above results, dimples are created within the lubricated area. In these cases, the film thickness at the inlet is the same as the minimum film thickness when the pad and moving surface are parallel. On the other hand, in the case of the Rayleigh step bearing, the step zone corresponds to an \u201copened dimple.\u201d The shape of the step zone could enter more lubricant into the lubricated area to develop a pressure distribution. Figure 12 shows a schematic of two arrangements of dimples. In one arrangement of dimples, dimples are created within the lubricated area. This arrangement of dimples is called \u201cenclosed dimple.\u201d In the other arrangement of dimples, the first dimple is created at the inlet side of the lubricated area and is therefore opened outside. This arrangement is called \u201copened dimple.\u201d When the dimples have lands at the leading side at Xdst\u00bc 0, the first dimple opens outside the lubricated area, as shown in Fig. 12. Figure 13 compares the variations in load W for various numbers of dimples between the enclosed-dimple case and openeddimple case at dimensionless dimple portion widths Ld of 0.2 and 0.4. For these results, the number of dimples is changed at a fixed dimensionless dimple portion width Ld. This means that the size of the dimple decreases with an increasing number of dimples. For Ld\u00bc 0.2, load W for the opened dimple is greater than that for the enclosed dimple. In the former case, the load is the greatest when there is only one dimple"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002133_icra.2014.6907617-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002133_icra.2014.6907617-Figure6-1.png",
        "caption": "Fig. 6. Artificially generated 6D rigid motion trajectory of a longish red object from a random starting configuration to the goal configuration inside the hole.",
        "texts": [
            " In this section we apply the GP over dual quaternions to an example setup incorporating large rotations and translations. The aim of this experiment is to learn a collision free 6D motion trajectory of a stylus from random starting configurations to a fixed goal configuration inside a hole. A set of 500 trajectories to sufficiently cover the 6D space is generated by the Open Motion Planning Library (OMPL) [14] using the path planner RRTconnect with path simplification and smoothing. Each trajectory consists of about 200 to 300 sequential poses visualized in Fig. 6. The GP over 6D rigid motions is trained with 20% randomly chosen sample poses from 108406 poses in total. The prediction is performed stepwise and we iteratively apply the predicted velocity to the current pose. To save computational effort, the prediction is based on the set of n \u2208 N closest poses according to the transformation magnitude measure (19). The starting configurations are restricted to the 6D space covered by the training trajectories. This assures positive correlation with training poses required by the GP prediction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003083_j.procs.2017.05.333-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003083_j.procs.2017.05.333-Figure4-1.png",
        "caption": "Fig. 4. Planetary gear transmission",
        "texts": [
            " In fact, whatever a gear-box has how many gears, in which all gears meshing can be classified as two kinds, the internal gear meshing and the external gear meshing. That is, a gear-box consists of several pairs of internal gear meshing, external gear meshing, or their combination. Jun Qiu et al. / Procedia Computer Science 109C (2017) 809\u2013816 813 Jun Qiu, Boxing Liu, Huimin Dong, Delun Wang / Procedia Computer Science 00 (2015) 000\u2013000 5 One of the most simple planetary gear transmissions, as shown in Fig. 4 (a), can be seen as a combination of internal gear meshing (Fig.4 (b)) and external gear meshing (Fig.4 (c)),Where C represents the planet carrier\uff0cP represents the planetary gear\uff0cS represents the sun gear\uff0cR represents the ring gear. Hence, a pair of external gears meshing is defined as an external gear transmission unit while a pair of internal gears meshing is called as an internal gear transmission unit. It is obvious that a planetary gear transmission can be viewed as the combination of the internal gear transmission unit and the external gear transmission unit. The gear transmission unit may be a single degree-of-freedom (DOF) mechanism (Fig. 4b) or a two DOFs mechanism (Fig.4c). They all have achieved the transmission of the kinematic, dynamic and structure characteristics information from the input shaft to the output shaft. The state characteristic of input shaft or output shaft is represented by the vector with sub-components. A gear-box has an input state vector denoted by Ri=(\u03c9i Mi ri)T and the output state vector Ro=(\u03c9o Mo ro)T. The relationship between the input state vector and output state vector can be expressed as o j iR A R  (1) where \u03bb\u03c9i, \u03bbMi, \u03bbri respectively are the transformation sub-matrix of rotate speed vector, torque vector and structure vector, they constitute the transformation matrix Aj"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003076_tmag.2017.2708748-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003076_tmag.2017.2708748-Figure1-1.png",
        "caption": "Fig. 1. Topologies. (a) 3-D structure of SMTLOM. (b) 3-D structure of SMTHLOM. (c) Cross-sectional view of SMTLOM. (d) Cross-sectional view of SMTHLOM.",
        "texts": [],
        "surrounding_texts": [
            "Novel Hybrid-Flux-Path Moving-Iron Linear Oscillatory Machine with Magnets on Stator Xiang Li1, Wei Xu1, Senior Member, IEEE, Caiyong Ye1, and Jianguo Zhu2, Senior Member, IEEE 1State Key Laboratory of Advanced Electromagnetic Engineering and Technology, School of Electrical and Electronic Engineering, Huazhong University of Science and Technology, Wuhan, 430074, China 2Faculty of Engineering and Information Technology, University of Technology Sydney, Sydney, 2007, Australia Whilst having various desired merits of high structural robustness and thus reliability, high thrust density, low fabrication cost, and so on, the stator-magnet moving-iron transverse-flux linear oscillatory machine (SMTLOM) suffers high consumption of rare earth magnetic material (high material cost) and heavy flux leakage at the stator outer region (underutilization of PMs). To mitigate the demerits, a novel stator-magnet moving-iron transversal-flux hybrid-flux-path linear oscillatory machine (SMTHLOM), with additional back iron and auxiliary radially-magnetized ferrite magnets, is proposed in this paper. The topology and operational principle of the proposed SMTHLOM is explained in details. The three-dimensional finite element analysis (3-D FEA) is introduced to investigate the sensitivities of main structural parameters, which enables the selection of optimal dimensions. Based on a set of comparison rules, a comprehensive comparison was conducted between these two topologies on various key indicators, such as the back electromotive force (EMF), thrust, and thermal distribution, etc. The simulation results show that the SMTHLOM has lower cost and smaller stator outer flux leakage, and thus is more suitable for some applications, e.g. fridge compressors, requiring low cost and high thrust density. Index Terms\u2014 Stator-magnet moving-iron transversal-flux hybrid-flux-path LOM (SMTHLOM), three-dimensional finite element analysis (3-D FEA), parameter sensitivity, thermal distribution, material cost. I. INTRODUCTION ERMANENT MAGNET (PM) LINEAR OSCILLATORY MACHINE (LOM) is promising while they are used to drive linear reciprocating devices, such as linear compressors, artificial hearts, breathing machines, and so on, for its high efficiency, power factor, and thrust density [1, 2]. To be compact, it usually adopts the tubular structure, whose flux path is kept in the same plane as the moving direction. Conventionally, its iron core consists of either circumferentially-laminated silicon steel sheets, or a solid entity made of soft magnetic composite (SMC) to reduce the iron loss [3]-[7]. The former would result in complex fabrication process while the latter suffer low strength and reliability. Traditionally, the PMs are placed on the mover and prone to irreversible demagnetization and physical damage due to the mechanical vibration. Because of the difficulty to protect the PMs, such structures always suffer high fabrication difficulty and low reliability [8]. A stator-magnet moving-iron transversal-flux LOM (SMTLOM) is proposed in [9], which uses a regular-laminated core that resembles the structure of a rotary switched reluctance machine and PMs mounted on the stator tooth surface. To certain extent, this could reduce the fabrication cost [10]. Higher reliability and lower fabrication cost can be achieved by inserting the PMs in the stator yoke as shown in Figs. 1 (a) and (c). This structure however still suffers high PM consumption (high material cost) and heavy flux leakage at the stator outer region (underutilization of PMs). To reduce the usage of rare earth magnets and flux leakage at stator outer region, this paper proposes a novel statormagnet moving-iron transversal-flux hybrid-flux-path LOM (SMTHLOM), which contains additional back irons and auxiliary radially-magnetized ferrite magnets around the stator outer region. In Section II, the new topology and its operational principle are explained in detail. Section III presents an investigation of the sensitivity of main structural parameters for an optimal scheme. Based on a set of comparison rules, i.e. the same size limitation, copper loss, and output thrust, various key indicators, such as the back electromotive force (EMF), thermal distribution, magnet consumption, and material cost, etc., are compared with those of the SMTLOM in Section IV. The comprehensive comparison results demonstrate that SMTHLOM has the advantages of lower cost and less flux leakage than SMTLOM."
        ]
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure9-1.png",
        "caption": "Fig. 9. Quantitative indices of the undercutting: a) radial and tangential; and b) tangential.",
        "texts": [
            " For the estimation of the cutting of the involute teeth the authors of the present paper propose two groups of quantitative indices to be defined. The first group of indices, defined as radial indices of the undercutting, specifies the degree of cutting of the involute teeth profile in a radial direction. Using the second group of indices, defined as tangential indices of the undercutting, the degree of the undercutting in a tangential direction of the teeth, is specified. In the presence of an undercutting of type I and type IIb, as it was already clarified, the section bq (Fig. 9a) of the involute tooth profile bqa is cut. In this case the location of the point q, obtained as a cross point of the gear fillet of the involute teeth profile, can be defined with the help of one of the following three possible parameters: the radius rq, at which the boundary point q is placed; the pressure angle \u03b1q of the involute curve in point q; the radius of the curvature \u03c1q of the involute curve in point q. The index radial undercutting (\u03b4r) defines immediately the length of the cut part of the involute tooth profile in a radial direction",
            " If the tooth undercutting is of the type I, for the calculation of the pressure angle \u03b1q, a correct numerical method or the below approximate formula [20] can be used tan\u03b1q \u00bc \u03c1q r b \u2248 m h a\u2212x\u22120:5z sin2\u03b1 r b : \u00f022\u00de In order to estimate what the length of the cut profile bq is, compared to the whole involute profile ba, the quantitative index a relative radial cutting (\u03bbr) is introduced. It is defined by the equation \u03bbr \u00bc \u03b4r ra\u2212rb \u22c5100 %\u00bd ; \u00f023\u00de ra is the radius of the addendum circle of the gear and the calculated value of \u03bbr is obtained in percentages. In the case of a non-undercut involute profile the radial indices of undercutting get a zero value (\u03b4r=0, \u03bbr=0%), and in the case of cutting of the whole segment ba of the involute profile, then \u03b4r=ra\u2212rb and \u03bbr=100%. On Fig. 9a it is seen that when the gear teeth are undercut in a radial direction, they are undercut simultaneously in a tangential direction, too. In the presence of an undercutting of type II\u0430 the teeth are cut only in a tangential direction (Fig. 9b), where the starting point b of the involute profile lies on the base circle of a radius rb. In order to determine the value of the tangential undercutting, from the center O of the gear a tangent line to the fillet curve fq is dropped, and the intersection u of the drawn tangent with the base circle, is determined. Using the arc \u2322bu, measured on the base circle, the quantitative index \u03b4t is defined where \u03b4t \u00bc \u2322bu \u00bc \u03b4 t m ; \u00f024\u00de a tangential undercutting, and the value \u03b4t\u204e appears as a coefficient of a tangential undercutting"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002286_j.triboint.2014.02.014-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002286_j.triboint.2014.02.014-Figure7-1.png",
        "caption": "Fig. 7. Model of piston secondary motion.",
        "texts": [
            " 6b) depends on the posture of the piston owing to the piston secondary motion. There are several possible postures of the piston in contact with the cylinder liner, such as one or more corners of the piston in contact with the cylinder liner. As shown in Fig. 6b, corner 1, corner 4, or both corners 1 and 4 to come into contact with the cylinder liner when the piston rotates in the counter-clockwise direction, whereas a clockwise rotation causes corner 2, corner 3, or both corners 2 and 3 to come into contact with the cylinder liner. As shown in Fig. 7, the impact force generated at the collision point of the piston and cylinder liner at corner 1 is represented by a hard stop, which consists of contact stiffness and damping that restrict the piston motion. Thus, the friction force due to contact of the piston skirt and cylinder liner at any instance can be determined based on the instantaneous piston position and orientation to determine the contact location of the piston skirt and cylinder liner. As shown in Fig. 7, the piston moves in the horizontal, vertical, and rotational directions where the motions are defined in y, x, and \u03b8 coordinates. The equation of motion of the piston system\u2014 which are the primary motion along the y-axis Fy, lateral motion along the x-axis Fx, and rotational motion about the piston pin axis M\u03b8 are shown below Fy \u00bcm\u20acyp \u00bc Frod cos \u03b8p Ff \u00f01\u00de Fx \u00bcm\u20acxp \u00bc Frod sin \u03b8p mg\u00fe \u2211 4 i \u00bc 1 Fx;i \u00f02\u00de M\u03b8 \u00bc \u2211 4 i \u00bc 1 \u00f0Fx;ili cos \u03b8\u00de mgly\u00fem\u20acxply m\u20acyplx Tf \u00f03\u00de Frodcos\u03b8p and Frodsin\u03b8p are the connecting rod force components in the x and y directions acting on the piston pin",
            " Ff is the total friction force measured by the dynamometer as discussed in the previous section. \u03a3Fx,i is the total impact force of piston corners 1\u20134 to the cylinder liner during piston slap. This value is calculated in Eq. (4) below. Tf is the total friction force torque about the piston pin axis, and li, lx and ly are the distance of each piston corner to the piston pin and the offset distances of the piston centre of gravity from the piston pin location in the x and y directions, respectively. Based on Fig. 7, the piston skirt comes into contact with the cylinder liner when the lateral displacement of any corner of the piston is greater than the clearance. The impact force generated by the piston skirt and moment of the piston are shown in Eq. (4); the acceleration, velocity, and displacement in both the lateral and rotational directions were measured by laser displacement sensors. Fx;i M\u03b8 \" # \u00bc mp 0 0 Ip \" # \u20acxp \u20ac\u03b8 \" # \u00fe ccp 0 0 c\u03b8 \" # _xp _\u03b8 \" # \u00fe kcp 0 0 k\u03b8 \" # xp \u03b8 \u00f04\u00de mp and Ip are the dynamic mass of a piston corner during piston slap and mass moment of the inertia of the piston, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001076_tcst.2017.2692732-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001076_tcst.2017.2692732-Figure3-1.png",
        "caption": "Fig. 3. Structure of a commercial ball-end arbor [55].",
        "texts": [
            " This kind of cutting processes is commonly used in various industries, such as dies manufacturing, aerospace, automotive, and biomedical industries. (For example, in biomedical industries, it is used to make small-size biomedical implants such as small bone implants, miniature left-ventricular assist devices, and finger joint replacements [53].) Recently, the cutting processes are moving toward higher precision and efficiency. To achieve a cutting process with an ultraprecision, high-speed ball-end arbors can be used [54] (see Fig. 3). During machining, heat flux is generated from different sources in three specific regions at the cutting point: the primary shear zone, the tool-chip interface, and the tool-workpiece interface [56]. It is known that more than 10% of the heat flux in the cutting edge goes into the cutting tool [57]. The heat generated during machining not only undesirably affects the dimension and form of the workpiece parts, but also reduces the useful life of the cutting tool [58]. To prevent these undesirable effects, control of the temperature in the cutting process can be taken into consideration [59]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003412_15397734.2018.1457446-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003412_15397734.2018.1457446-Figure2-1.png",
        "caption": "Figure 2. 3D FEM model of axle-wheel-rail set, (a) shows the whole model and (b) the position of the wheel flat. The leading edge is at the front of the wheel flat in the motion direction.",
        "texts": [
            " In this study,ANSYS APDL softwareis used as preprocessing, processing, and post- processing software. The wheel and the rail profilesarethe standard UIC920 mm freight wheeland UIC60, respectively. These profiles are first modeled as 2D surfaces according to Figure 1. To obtain more accurate results, smaller elements are used in the desired area with the smallest dimension of 1.87mm\u00d71mm. The 3D model is obtained from these 2D elements by revolving or extruding.Note that, the 2D elements are merely used for modeling purpose and are cleared in the end. Figure 2(a) illustrates the 3D finite element model of the axle-wheel-rail set. The wheel starts rolling at an initial velocity and reaches the desired region, a 200mm halfway part of the rail between two supporting sleepers, after a full rotation. The wheel flat is shown in Figure 2(b). The leading edge is at the front of the wheel flat in the motion direction as shown in Figure 2(b). A 3D 8-node structural elementis employed for modeling of the wheel-rail structure.The contact type atthewheel-rail interface issurface-to-surface. A standard unilateral contact is used as the contact surface behavior in which, the normal pressure is zero when separation occurs (ANSYS, 2013). A coefficient of friction equal to 0.5is assumed (Wei et al. 2016).This high value guarantees the rollingmotion of the wheel.It is found from the present analyses thatthetraction coefficients arein fact much less than 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002250_s12206-014-0919-3-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002250_s12206-014-0919-3-Figure1-1.png",
        "caption": "Fig. 1. Actions of the O-ring seal and X-ring seal.",
        "texts": [
            " Although the O-ring seal offers a reasonable approach to the ideal hydraulic seal, they are not considered the best solution to all sealing problems. Various studies have reported that there are certain limitations on the use of O-ring seals, including high temperature applications and high rubbing speeds. Oring seals are not recommended for speeds of less than 0.005 m/s when the pressure difference is less than 2.75 MPa. Although an O-ring under pressure slightly prevents spiral rolling as shown in Fig. 1 [14], this situation is not adequate. A good solution to avoid spiral failure is the use of Quad-rings or X-rings. X-rings are used in many dynamic applications in which O-rings provide unsatisfactory performance. An X-ring is a four-lobed seal designed for improved seal lubrication and to prevent rolling of the seal, or spiral failure, as shown in Fig. 1 [15]. Various parameters are used to define the geometry of the X-ring, unlike other sealing rubber elements. Eq. (1) gives the relationship between the three basic geometric parameters of an X-ring [16]: 2( )H L r  , (1a)  2 2 ( ) 2 1R H r   , (1b)  2 ( ) ( 2 ) R r L R r R r R     , (1c) where the variables are defined in Fig. 2(a). Eq. (1) is adequate to describe the geometry of the X-ring when the desired cross-section that best fits the groove is known. The stress components using Muskhelishvilli's complex function and the Airy stress function [17] are given by      Re 2x z z z z        ,      Re 2y z z z z        , (2)    Imxy z z z      , where for convenience, we have put ( ) ( ),z z  ( ) ( )z z  and z x iy  "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000460_s12239-015-0043-0-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000460_s12239-015-0043-0-Figure1-1.png",
        "caption": "Figure 1. External meshing gear pairs schematic.",
        "texts": [
            " Besides, this paper proposes a configuration to improve the transmission efficiency of the gearbox after analyzing simulation results. Gear meshing power losses can be divided into two parts, sliding and rolling power losses. Under the condition of EHL, the accuracy of mathematic model is mainly determined by friction coefficient, load distribution, location of meshing point and oil film thickness. In this paper, the research on these factors is conducted and the average gear meshing mechanical power losses within a meshing cycle is deduced. 2.1. Spur Gears Figure 1 shows the external meshing gear pairs schematic. B1, B2 : start of mesh cycle and end of mesh cycle, respectively : actual length of line of action, mm C, D : start of single-tooth contact and end of singletooth contact, respectively : theoretical length of line of action, mm O : meshing point P : pitch point e, e1, e2 : length of , and , mm As shown in Figure 1, mesh cycle is defined as the motion from the engaging-in point to the engaging-out point of one pair of meshing teeth. The line of action is divided into four parts (PC, CB1, PD, DB2) to be used in the calculation. Take pitch point P as the original point. Firstly, the transient meshing power losses at the meshing point are calculated. It is a function of location of meshing point. And then the integrations along the direction of engagingin and engaging-out are calculated respectively. Eventually, the average power losses in the whole mesh cycle are obtained",
            " As the contact condition of the engaged surfaces of the pair of gears is a successive one, according to the knowledge of Contact Mechanics (Johnson, 1992), the normal velocity components on the meshing point of the two engaged gears are equal. 2.1.1. Instantaneous sliding velocity of meshing point When meshing point O is inside (i.e. ), the sliding velocity of pinion on the meshing point is larger than that of gear; When meshing point O is inside (i.e. ), the sliding velocity of pinion on the meshing point is smaller than that of gear, which is also easy judged from Figure 1. Therefore the instantaneous sliding velocity of meshing point can be written as: (1) where \u03c91 is the rotational angular speed of pinion, and \u03c92 the rotational angular speed of gear. By trigonometric function and property of meshed involute gears, it is easy to deduce the following expressions: (2) (3) According above formula, we can draw the following equation: (4) where n1 is the rotational speed of pinion, i the gear ratio, and S the length of . 2.1.2. Instantaneous rolling velocity of meshing point Rolling is defined as the relative angular velocity of the bodies rotating around the axis in tangent plane (Johnson, 1992)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001515_aad52b-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001515_aad52b-Figure1-1.png",
        "caption": "Figure 1. Schematic diagram of experimental equipment.",
        "texts": [
            " The stability of the solid solution phase in HAEs is closely related to the valence electron concentration (VEC), so the values of VEC are also calculated and discussed in this paper. AlCoCrFeNiSix HEAs with different Si contents (x = 0, 0.5, 1.0, 1.5 and 2.0) were synthesized by arc melting, and were composed of Al, Co, Cr, Fe, and Ni (purity\uff1e99.9%) mixed with various concentrations of Si (purity\uff1e99.9%) under an atmosphere of argon. When x = 0, 0.5, 1.0, 1.5 and 2.0, the alloys were denoted as Si 0, Si 0.5, Si 1.0, Si 1.5 and Si 2.0, respectively. Figure 1 shows a schematic diagram of the experimental equipment. The ingots were melted more than four times to achieve a chemical homogeneity, followed by suction casting into copper molds. A field-emission scanning electronic microscope (FESEM) (SU8010, Hitachi, Japan), equipped with an energy dispersive spectrometer (EDS) and an electron probe microanalyzer (EPMA-1720) were used to investigate chemical compositions of different phases of the HEAs. Xray diffraction (XRD) (Shimadzu 7000, Kyoto, Japan) was used to identify the constituent phases with the 2\u03b8 scan ranging from 20\u00b0 to 100\u00b0 at a speed of 6\u00b0/min and Cu-K\u03b1 radiation (\u03bb = 0",
            " As shown in Table 3, it is clear seen that the radius of Si is the 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 A cc ep te d nu Journal XX (XXXX) XXXXXX Author et al 6 smallest among all elements, leading to an increase in the atomic size difference with the addition of Si content. Moreover, the formation of the hard Cr3Si phase also contributes to an increase in the hardness of the alloy. In addition, the phase evolution of the alloy also improved hardness. Since the phase transforms from a FCC and BCC1 to a BCC2 structure, it is clear that the BCC structure has a higher hardness than the FCC structure [21]. The compressive stress-strain curve of an as-cast AlCoCrFeNiSix HEAs alloy is shown in Fig.1. It is found that the strength of HEAs increased significantly with the addition of Si content, while the alloy ductility decreased. As the Si content increases from 0.5 to 2.0, the strength of the alloy continues to increase significantly by the sacrifice of ductility. The plastic deformation of alloys without Si obtained more than 25%. However, the plastic deformation was less than 15% with the increase of Si content from 0.5 to 2.0. The increase in strength may be attributed to the solid solution of Si, which enhances the difference in atomic size"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003140_ffe.12681-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003140_ffe.12681-Figure2-1.png",
        "caption": "FIGURE 2 MD simulation model [Colour figure can be viewed at wileyonlinelibrary.com]",
        "texts": [
            " With an aim to identify the characteristics and behaviors of the crack initiation and propagation under dynamic maximum tensile stress at the spur gear tooth root, the MD simulation is carried out. In general, the essential component of the spur gear material is ferrite which has bcc crystal structure and a worse ability in dissolving carbon. If carbon atoms are introduced in the MD simulation, the analysis of the mechanisms of crack initiation and propagation will become more complex. To simplify the analysis, this work focuses on the pure bcc iron by neglecting the alloy elements. tooth. B, Strain rate at the upper area of the tooth root In Figure 2, the MD simulation model is developed by the software LAMMPS,23 and the crystal directions of the X, Y, and Z axes are [0 1 0], [1 0 1], and 1 0 1 , respectively. The dimensional size of the model is 120a0 \u00d7 60a0 \u00d7 6a0, where a0 is the crystal constant of the bcc iron and its value is 2.8553 \u00c5. As detailed in Section 1, the number of surface atoms is very large if the free boundaries are introduced, and for the small model in this MD simulation, it means a high surface area to volume ratio, which does not coincide with the considered macroscopic system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003379_0954410018764472-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003379_0954410018764472-Figure1-1.png",
        "caption": "Figure 1. The morphing aircraft with variable sweep angle and variable span.",
        "texts": [
            " if and only if there exists a full-block multiplier such that M11 M12 I 0 M21 M22 2 4 3 5T diag Y ,Q n o M11 M12 I 0 M21 M22 2 4 3 55 0 and for any 2 ? I diag\u00f0 \u00de TY I diag\u00f0 \u00de 50 In this section, the longitudinal nonlinear dynamic equations, as the key component in the design process, are derived based on Kane multi-body method at first. Then the polynomial parameter-dependent LPV system is developed for the following controller synthesis. Nonlinear dynamic modeling of the morphing aircraft Here, we consider a kind of large-scale morphing aircraft,25 depicted in Figure 1, in which their airfoils can be driven by smart actuators to produce a certain elongation or contraction in wing span and rotation in wing sweep angle. The geometry of the fixed parts (fuselage, vertical tail, horizontal tail, etc.) of the morphing aircraft is obtained according to The Teledyne Ryan BQM-34 \u2018\u2018Firebee\u2019\u2019. The airfoils on both sides, as the morphing subjects, are consisted of inner wings and outer wings. In wing morphing process, the inner can rotate with Ob and the outer will keep level and move linearly along with the airfoil"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000056_j.precisioneng.2014.11.008-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000056_j.precisioneng.2014.11.008-Figure1-1.png",
        "caption": "Fig. 1. The overall mechanism and",
        "texts": [
            " Design requirements The main requirements on the automatic variable preload system are the following three characteristics: (1) enough elasticity in order to ensure reversible operation of the system, (2) appropriate mass distribution in order to generate the required amount of centrifugal force, and (3) a simple mechanism for converting the radial centrifugal force to axial preload so that by raising the centrifugal force, axial preload is decreased. 2.2. Detailed design The overall mechanism of variable preload system is shown in Fig. 1. It consists of a specially designed ring which acts as a spring (hereinafter called spring) equipped with centrifugal mass elements, a spacer ring used for transmitting the force, and a locknut. The components of the variable preload system are assembled on the spindle and rotate at spindle speed. The spring is made of a ring with radial slits that divide it into equally-spaced sections. Each of these sections behaves like a cantilever beam, and thus provides considerable elastic behavior allowing the piece to act as an axial spring"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000683_icra.2013.6631258-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000683_icra.2013.6631258-Figure6-1.png",
        "caption": "Fig. 6. Shapes of manipulator : which have (a) peak of 1w2 in Fig. 7(a) with DRME and (b) peak of 2nd-link RMM in Fig. 7(b) with RME.",
        "texts": [
            "0 10.0 Lower body 0.1 2.0 10.0 Upper arm 0.31 2.3 0.03 Lower arm 0.24 1.4 1.0 Hand 0.18 0.4 2.0 Waist 0.27 2.0 10.0 Upper leg 0.38 7.3 10.0 Lower leg 0.40 3.4 10.0 Foot 0.07 1.1 10.0 Total 1.7 63.8 Length, mass and coefficient of viscous friction of each link and joint are set to be 0.3[m], 1.0[kg], and 2.0[N\u00b7m\u00b7s/rad]. In this simulation, we assume that tip of link-2 and link-4 are always placed y = 0, that is, when q2 and q4 are given, q1 and q3 are set as q1 = \u2212q2/2.0 and q3 = \u2212(q2 +q4)/2.0. Figure 6(a) and (b) depict the DRME (scaled) and RME (scaled) with manipulator shapes, which indicate the peak of 2nd-link-DRMM distribution and 2nd-link reconfiguration manipulability measure (RMM) distribution shown in Fig. 7. In Fig. 7, the peak of 2nd-link RMM at q2 = 90\u25e6 and q4 = 90\u25e6. On the other hand, the peak of 2nd-link DRMM at q2 = 118\u25e6 and q4 = 141\u25e6. Difference of the results is caused from considering dynamical peculiarities or not. We discuss a biped robot whose definition is depicted in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure4.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure4.1-1.png",
        "caption": "Fig. 4.1 Simplified scheme of the two-pole induction machine magnetic circuit",
        "texts": [
            " So, the magnetizing reactance xm can be expressed through the values of the magnetizing reactance values of the electric machine magnetic circuit regions. Below, we consider the features of the determination of the magnetizing reactance values of the magnetic circuit regions. In this chapter for this purpose, we use the total current law. The basic construction elements of the active zone of an electric machine (the air gap, slot wedge regions, slot regions occupied by the windings, stator and rotor yoke regions) form the magnetic circuit for the magnetizing flux\u0424m (Fig. 4.1). The construction elements of the active zone of an electric machine are considered as the regions of its magnetic circuit. When calculating the electric machine magnetic circuit, the following basic simplifying assumptions are usually used: in the magnetic circuit regions, the magnetic flux has the same value equal to\u0424m; through the air gap and the stator and rotor teeth regions, magnetic flux passes in the radial direction, whereas in the stator and rotor yoke regions the flux is directed tangentially; in the axial direction, the stator and rotor iron is distributed continuously; the presence of the ventilation channels, insulation spaces and finite lengths of the stator and rotor iron are taken into account by using the special factors in the final expressions; the air gap magnetic induction determining the value of the flux \u0424m is given on the surface passing through the middle of the air gap, and it is sinusoidal within the pole pitch",
            " The presence of the flattening of the magnetic induction curve is taken into account by the special factor; the stator and rotor surfaces are assumed \u00a9 Springer International Publishing Switzerland 2015 V. Asanbayev, Alternating Current Multi-Circuit Electric Machines, DOI 10.1007/978-3-319-10109-5_4 95 as smooth; and the presence of the slots on the stator and rotor surfaces is taken into account by the calculated value of the air gap length. The magnetic circuit of an electric machine is represented as the line located in its cross section and passing in the middle of the magnetic pole formed by the winding (Fig. 4.1). Then the quite complicated task of determining the magnetic field distribution is reduced to the more simple and tangible task of the magnetic circuit calculation. Now for the calculation of the magnetic circuit, the total current law can be used 1 2 \u00fe L Hdl \u00bc Fm \u00f04:1\u00de where H is the magnetic field strength, and dl is the elementary part of the contour L (Fig. 4.1). lyas aS 96 4 The Magnetic Circuit Regions: The Magnetizing Reactance Values When calculating the magnetic circuit by equation (4.1) instead of Fm, the amplitude value of the stator winding mmf defined on the pole and phase is used Fm \u00bc m1w1kw1 p\u03c0 ffiffiffi 2 p Im \u00f04:2\u00de In (4.2), the magnitude of the current Im represents the effective value. In practical calculations, it can be assumed that the contour of the passing of the magnetic flux L (Fig. 4.1) can be broken down into the discrete regions, within which the magnetic field strength H takes a constant value. Then from (4.1) and (4.2), we can receive 1 2 \u00fe L Hdl \u00bc 1 2 Xm n\u00bc1 Hnlyn \u00bc m1w1kw1 p\u03c0 ffiffiffi 2 p Im \u00f04:3\u00de where Hn and lyn are the magnetic field strength and length of the nth magnetic circuit region, and m is the total number of regions on which the magnetic circuit is subdivided. For the magnetic field strength of the nth region, we have Hn \u00bc Bn \u03bcn \u00bc \u03a6m \u03bcnSn \u00bc \u03a6m \u03bcn 2=\u03c0\u00f0 \u00de\u03be0\u03c4nlan \u00bc \u03a6m \u03bcnbnlan \u00f04:4a\u00de where Bn\u00bc\u0424m/Sn; Sn\u00bc (2/\u03c0)\u03be0\u03c4nlan\u00bc bnlan; bn\u00bc (2/\u03c0)\u03be0\u03c4n;Bn, Sn and \u03bcn are the magnetic induction, cross section, and magnetic permeability of the nth region; bn and lan are the calculated values of the width and axial length of the nth magnetic circuit region; and n\u00bc zs, ks, 0, kR, zR",
            "3) can be written as EmA \u00bc 1 \u03c0 l0 \u03c91\u03beB\u03c40 \"Xk n\u00bc1 lyn \u03bcnSn \u00fe Xm n\u00bc m k\u00f0 \u00de lyn 2\u03bcnSn # ImA \u00bc xmAImA \u00f04:6\u00de where xmA \u00bc 1 \u03c0 l0 \u03c91\u03beB\u03c40 \"Xk n\u00bc1 lyn \u03bcnSn \u00fe Xm n\u00bc m k\u00f0 \u00de lyn 2\u03bcnSn # Here, for the calculated value of the active length of the air gap l0, we have l0 \u00bc l1 nBbB \u00fe 2\u03b4 where l1 is the total stator length, and nB and bB are the number and width of the ventilation channels. For electric machines with a small air gap (\u03b4 bB), it follows that l0\u00bc l1 nBbB. With a large air gap (\u03b4 bB), we have that l0\u00bc l1 + 2\u03b4. In (4.6) the pole pitch \u03c40 is determined on the surface passing through the middle of the air gap. In (4.6), the value of xmA represents the magnetizing reactance of an electric machine. Here, it is expressed in the specific system of units. The magnetic circuit shown in Fig. 4.1 can be represented as consisting of the seven regions (m\u00bc 7) corresponding to the stator joke (as), stator slot area occupied by the phase winding (zs), stator slot wedges (ks), air gap (0), rotor slot wedges (kR), rotor slot area occupied by the winding (zR), and the rotor joke (aR). Then the expression (4.6) for the reactance xmA can be written in the form 98 4 The Magnetic Circuit Regions: The Magnetizing Reactance Values xmA \u00bc 1 \u03c0 l0 \u03c91\u03beB\u03c40 \"Xk n\u00bc1 lyn \u03bcnSn \u00fe Xm n\u00bc m k\u00f0 \u00de lyn 2\u03bcnSn # \u00bc 1 1 x\u03c4as \u00fe 1 xzs \u00fe 1 xks \u00fe 1 x0 \u00fe 1 xkR \u00fe 1 xzR \u00fe 1 x\u03c4aR \u00f04:7\u00de where xn \u00bc \u03c91\u03beB\u03bcn\u03c40 \u03c0 Sn l0lyn n \u00bc zs, ks, 0, kR, zR\u00f0 \u00de and x\u03c4n \u00bc 2\u03c91\u03beB\u03bcn\u03c40 \u03c0 Sn l0lyn n \u00bc as, aR\u00f0 \u00de",
            " From these expressions it follows that to obtain the analytical expressions for the considered magnetizing reactance values, it is necessary to have the linear dimensions of the magnetic circuit regions bn (across the magnetic flux lines), and lyn (in the direction of the magnetic flux lines), and also the values of the magnetic permeabilities \u03bcn and the axial lengths lan of the magnetic circuit regions. The linear dimensions bn (except for the stator and rotor joke regions) reflect the value of the calculated length of the pole arc. In the case of the stator and rotor joke regions for the values of bn, we accept their heights. The linear dimensions lyn are determined (except for the stator and rotor joke regions) on the basis of Fig. 4.1. According to Fig. 4.1, the value of lyn represents the double lengths of the corresponding magnetic circuit regions. The axial length of these regions lan is determined by taking into account the features of their design. The stator equivalent circuit in Fig. 4.4 is determined by the reactance values x01, xks, xzs, and x\u03c4as. The values of x01, xks, xzs, and x\u03c4as reflect the magnetizing reactance values of the corresponding stator magnetic circuit regions. Below, we will define these reactance values. The Magnetizing Reactance of the Air Gap As applied to the air gap, the expression (4",
            "35) corresponds to the average value of the magnetic induction acting in the radial direction in the stator slot wedge region within the teeth division. Therefore in relation to this radial magnetic induction, the considered stator tooth region can be represented as the conditional uniform layer characterized by the average value of the radial component of the magnetic permeability equal to \u03bcyks. The calculated pole arc on the surface of this stator tooth region is bks \u00bc 2=\u03c0\u00f0 \u00de\u03be0\u03c4ks \u00f04:36\u00de where \u03c4ks is the pole pitch on the stator bore surface. According to Fig. 4.1, the double thickness of the stator slot wedge region is lyks\u00bc 2hks. Now using the conditions lyks\u00bc 2hks and bks\u00bc (2/\u03c0)\u03be0\u03c4ks, we obtain for the factor \u03bdks 4.3 The Stator Magnetic Circuit Regions: The Magnetizing Reactance Values 107 \u03bdks \u00bc \u03beLksbks lyks \u00bc \u03beLks\u03be0\u03c4ks \u03c0 hks \u00f04:37\u00de By the expressions (4.28), (4.33), and (4.37), the reactance xks takes xks \u00bc \u03c91\u03beB\u03be0\u03bcyks\u03c4 2 ks \u03c02hks \u03beLks \u03c4ks=\u03c40 \u00bc \u03c91\u03beB\u03be0\u03bcyks\u03c4 2 ks \u03c02hks \u03beLks \u03be\u03c4ks \u00f04:38\u00de where \u03be\u03c4ks\u00bc \u03c4ks/\u03c40. Considering that t01\u00bc (2p\u03c40/Z1), the expression (4",
            " In this case, this expression should be written for the geometry of the considered wound part of the stator tooth region. In the expressions (4.42) and (4.43), \u03bcyzs reflects the average value of the radial component of the magnetic permeability, in relation to which the considered wound part of the stator tooth region can be represented as the conditional uniform layer. The calculated pole arc on the surface of this stator region is bzs \u00bc 2=\u03c0\u00f0 \u00de\u03be0\u03c4zs \u00f04:44\u00de where \u03c4zs is the pole pitch on the inner surface of the wound part of the stator tooth region. According to Fig. 4.1, the double length of the wound part of the stator tooth region is lyzs\u00bc 2hzs. Now using the conditions lyzs\u00bc 2hzs and bzs\u00bc (2/\u03c0)\u03be0\u03c4zs, we can receive for the factor \u03bdzs \u03bdzs \u00bc \u03beLzsbzs lyzs \u00bc \u03beLzs\u03be0\u03c4zs \u03c0 hzs \u00f04:45\u00de The magnetizing reactance of the wound part of the stator tooth region takes from (4.41) and (4.45) xzs \u00bc \u03c91\u03beB\u03be0\u03bcyzs\u03c4 2 cz \u03c02hzs \u03beLzs \u03c4zs=\u03c40 \u00bc \u03c91\u03beB\u03be0\u03bcyzs\u03c4 2 cz \u03c02hzs \u03beLzs \u03be\u03c4zs \u00f04:46\u00de where \u03be\u03c4zs\u00bc \u03c4zs/\u03c40. Taking into account that t01\u00bc (2p\u03c40/Z1), the expression (4.46) can be written in another form xzs \u00bc \u03c91\u03beB\u03be0\u03bcyzs\u03c4 2 cz \u03c02hzs \u03beLzs \u03be\u03c4zs \u00bc \u03c91\u03bcyzs \u03beB\u03be0\u03c4cz\u03c40\u03beLzs \u03c02hzs \u00bc \u03c91\u03bcyzst01 \u03beB\u03be0\u03c4zsZ1\u03beLzs 2p\u03c02hzs \u00bc \u03c91\u03bcyzst01\u03bbzs\u03beLzs \u00bc \u03c91\u03bcyzst01\u03bb 0 zs \u00f04:47\u00de where \u03bb 0 zs \u00bc \u03bbzs\u03beLzs; \u03bbzs \u00bc \u03beB\u03be0\u03c4zsZ1 2p\u03c02hzs is the permeance factor for the flux \u0424m in the wound part of the stator tooth region",
            "46) the reactance xzs in the phase system of units xzs \u00bc xzsAk1A \u00bc \u03c91\u03bcyzst01\u03bb 0 zsk1A \u00bc \u03c91\u03bcyzst01\u03bb 0 zs 2m1 w1kw1\u00f0 \u00de2l0 p\u03c40 \u00bc \u03c91\u03bcyzsl0\u03bb 0 zs 4m1 w1kw1\u00f0 \u00de2 Z1 \u00bc \u03c91\u03bcyzsl0\u03bb 0 zsk 2 1 \u00f04:48\u00de The Magnetizing Reactance of the Stator Joke Region From the expression (4.11), we can obtain for this reactance x\u03c4as \u00bc 2\u03c91\u03beB\u03bcas\u03c40 \u03c0 \u03bdas \u00f04:49\u00de where \u03bdas\u00bc (\u03beLasbas/lyas); \u03beLas\u00bc (laas/l0); laas is the axial length of the stator joke iron. In accordance with (4.49), it is necessary to find the values of bas, lyas, \u03bcas and laas. From Fig. 4.1 it follows that for the stator joke region, the linear dimension bas is determined from the condition bas\u00bc has, where has is the height of the stator joke region. In (4.49), the value of lyas reflects the conditional length of the magnetic line on which the average value of the magnetic field strength (magnetic field strength on the neutral between the stator polesHasH) provides the necessary mmf to conduct the flux\u0424m in the stator joke region. On the other hand, the value of mmf (necessary to conduct the flux \u0424m through the stator joke region) can be determined by integrating the magnetic field strength along the part of the joke surface corresponding to the length of one pole pitch",
            "91) is determined by the expression of the form (4.34). In this case, this expression should be written for the geometry of the rotor slot wedge region. In the expressions (4.90) and (4.91), \u03bcykR reflects the average value of the radial component of the magnetic permeability, in relation to which the considered rotor slot wedge region can be presented as the conditional uniform layer. The calculated pole arc on the rotor surface is bkR \u00bc 2=\u03c0\u00f0 \u00de\u03be0\u03c4kR \u00f04:92\u00de where \u03c4kR is the pole pitch on the rotor surface. According to Fig. 4.1, the double length of the rotor slot wedge region is lykR\u00bc 2hkR. Now using the conditions lykR\u00bc 2hkR and bkR\u00bc (2/\u03c0)\u03be0\u03c4kR, we can receive for the factor \u03bdkR \u03bdkR \u00bc \u03beLkRbkR lykR \u00bc \u03beLkR\u03be0\u03c4kR \u03c0 hkR \u00f04:93\u00de From (4.89) and (4.93), the magnetizing reactance of the rotor slot wedge region takes xkR \u00bc \u03c91\u03beB\u03be0\u03bcykR\u03c4 2 kR \u03c02hkR \u03beLkR \u03c4kR=\u03c40 \u00bc \u03c91\u03beB\u03be0\u03bcykR\u03c4 2 kR \u03c02hkR \u03beLkR \u03be\u03c4kR \u00f04:94\u00de where \u03be\u03c4kR\u00bc \u03c4kR/\u03c40. The expression (4.94) can be presented in another form. For this purpose, we use expression (4.86) and the condition \u03c40\u00bc (Z2t02/2p)",
            " 120 4 The Magnetic Circuit Regions: The Magnetizing Reactance Values In the expressions (4.98) and (4.99), \u03bcyzR reflects the average value of the radial component of the magnetic permeability, in relation to which the wound part of the rotor tooth region can be represented as the conditional uniform layer. The calculated pole arc on the outer surface of this rotor region is bzR \u00bc 2=\u03c0\u00f0 \u00de\u03be0\u03c4zR \u00f04:100\u00de where \u03c4zR is the pole pitch on the outer surface of the wound part of the rotor tooth region. According to Fig. 4.1, the double length of the wound part of the rotor tooth region is lyzR\u00bc 2hzR. Now using the conditions lyzR\u00bc 2hzR and bzR\u00bc (2/\u03c0)\u03be0\u03c4zR, we can receive for the factor \u03bdzR \u03bdzR \u00bc \u03beLzRbzR lyzR \u00bc \u03beLzR\u03be0\u03c4zR \u03c0 hzR \u00f04:101\u00de From (4.97) and (4.101), the magnetizing reactance of the wound part of the rotor tooth region obtains xzR \u00bc \u03c91\u03beB\u03be0\u03bcyzR\u03c4 2 zR \u03c02hzR \u03beLzR \u03c4zR=\u03c40 \u00bc \u03c91\u03beB\u03be0\u03bcyzR\u03c4 2 zR \u03c02hzR \u03beLzR \u03be\u03c4zR \u00f04:102\u00de where \u03be\u03c4zR\u00bc \u03c4zR/\u03c40. By the expression (4.86), we can have for the reactance xzR shown in (4.102) xzR \u00bc \u03c91\u03beB\u03be0\u03bcyzR\u03c4 2 zR \u03c02hzR \u03beLzR \u03be\u03c4zR \u00bc \u03c91\u03bcyzR \u03beB\u03be0\u03c4zR\u03c40\u03beLzR \u03c02hzR \u00bc \u03c91\u03bcyzRt02 \u03beB\u03be0\u03c4zRZ2\u03beLzR 2p\u03c02hzR \u00bc \u03c91\u03bcyzRt02\u03bbzR\u03beLzR \u00bc \u03c91\u03bcyzRt02\u03bb 0 zR \u00f04:103\u00de where \u03bb 0 zR \u00bc \u03bbzR\u03beLzR; \u03bbzR \u00bc \u03beB\u03be0\u03c4zRZ2 2p\u03c02hzR is the permeance factor for the flux \u0424m in the wound part of the rotor tooth region",
            "4 The Rotor Magnetic Circuit Regions: The Magnetizing Reactance Values 121 x\u03c4aR \u00bc 2\u03c91\u03beB\u03bcaR\u03c40 \u03c0 \u03bdaR \u00f04:105\u00de where \u03bdaR\u00bc (\u03beLaRbaR/lyaR); \u03beLaR\u00bc (laaR/l0); laaR is the axial length of the iron of the rotor joke region. In accordance with (4.105), it is necessary to find the values of baR, lyaR, \u03bcaR and laaR. The magnetic permeability of the rotor joke region \u03bcaR is determined by the average value of the magnetic induction on the neutral between the rotor poles. The magnetic induction on the neutral between the rotor poles represents the tangential magnetic field. Therefore we can write that \u03bcaR\u00bc \u03bcxaR. From Fig. 4.1 it follows that for the rotor joke region, the linear dimension baR is determined from the condition baR\u00bc haR, where haR is the height of the rotor joke region. To calculate the conditional length of the average magnetic line in the rotor joke region lyaR, we can use the expressions (4.63) and (4.76). In the planar coordinate system, lyaR can be determined by the expression (4.63), if in it the value of \u03b2as to reduce by \u03b2aR\u00bc \u03c0/ \u03c4aR (since k\u03bcaR\u00bc 1.0) and the stator joke region height has by haR. Then, we can obtain lyaR \u00bc 2 \u03c4aR \u03be0\u03c0 \u03b2aRhaR th\u03b2aRhaR \u00f04:106\u00de The conditional length of the average magnetic line in the cylindrical rotor joke region can be calculated by the expression (4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000748_tie.2013.2276025-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000748_tie.2013.2276025-Figure1-1.png",
        "caption": "Fig. 1. Induction heating of rotating cylindrical billet in static magnetic field produced by direct-current carrying coil.",
        "texts": [
            " Dole\u017eel is with the University of West Bohemia, 306 14 Pilsen, Czech Republic, and with Czech Technical University, 166 27 Prague, Czech Republic and also with the Institute of Thermomechanics, Academy of Sciences of the Czech Republic, 182 00 Prague, Czech Republic (e-mail: idolezel@kte.zcu.cz). Digital Object Identifier 10.1109/TIE.2013.2276025 drawbacks of this way of processing can be found in numerous references (see, e.g., [9]\u2013[12]). Much recent (about ten years) is the latter technology. The schematic arrangement of the process is shown in Fig. 1. However, even in this paper, several serious issues have to be addressed [13]\u2013[18]. The crucial one is the generation of a sufficiently high magnetic field and, consequently, eddy currents and Joule losses in the billet. The only solution is using very high field currents, which requires the presence of a superconducting system with all necessary infrastructures. Now, the Joule losses in the windings are not decisive. Certain part of the energy delivered to the system, however, must be used for cooling and covering the mechanical losses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000343_s40430-013-0053-7-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000343_s40430-013-0053-7-Figure1-1.png",
        "caption": "Fig. 1 Physical configuration of the slider bearing",
        "texts": [
            " [16] and this theory has been successfully used to study the hydrodynamic lubrication of rough slider bearings with non-Newtonian fluids by various researchers [17, 18]. In this paper, an attempt has been made to analyze the effect of micropolarity\u2013roughness interaction on the performance of hydrodynamic lubrication of slider bearings of various shapes viz; plane slider, exponential, hyperbolic and secant. The probability density function for the random variable characterizing the surface roughness is assumed to be asymmetrical with non-zero mean. Figure 1 shows the schematic diagram of the bearing configuration under study. It consists of two surfaces separated by a lubricant film. The lower surface of the bearing is moving with a constant velocity U in its own plane while the upper surface is at rest. To represent the surface roughness, the mathematical expression for the film thickness is considered to be consisting of two parts. H\u00f0x\u00de \u00bc h\u00f0x\u00de \u00fe hs \u00f01\u00de where h(x) is the mean film thickness and hs is a randomly varying quantity measured from the mean level and thus characterizes the surface roughness and L1 being the length of the bearing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001022_tia.2017.2683439-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001022_tia.2017.2683439-Figure9-1.png",
        "caption": "Fig. 9. No load magnetic field distribution for the motor without damaged regions at rated voltage.",
        "texts": [
            " Looking at this table, it is immediate observing the negative effect of the punching process, both in term of magnetizing field and losses increases. As aforementioned, since the motor prototypes do not have the rotor cage, a dedicated V/f law has been used in the simulations to synchronize the rotor at the desired speed. Then, the motors were analyzed at the no-load conditions with the imposed sinusoidal supply voltage, determining the flux density distributions. An example of the field density distribution is presented in Fig.9, where the \u2018A\u2019 lines separate the regions which may have different magnetic properties, as in the case of the non-annealed motor model. The direction of the magnetic flux plays an important role. In the computation of the iron losses in the damaged zones, As discussed in [11], the losses inside the damaged area change if the magnetic flux flows parallel to the \u2018green\u2019 material with respect to the situation in which the magnetic flux flows perpendicularly from the damaged area to the undamaged one"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001814_gt2015-43935-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001814_gt2015-43935-Figure2-1.png",
        "caption": "Figure 2. Support structures (red) used for the (a) horizontal, (b) 45\u00b0, and (c) vertical build directions.",
        "texts": [
            " To examine the build direction effects on surface roughness, geometric tolerance, and ultimately the pressure loss and heat transfer, different test coupons were designed and manufactured in-house. Shown in Figure 1, the overall size of the coupons were 2 Copyright \u00a9 2015 by ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 01/24/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use 25.4mm long x 25.4mm wide x 3.048mm high. Each coupon had 15 channels spaced 2.5 diameters apart, with a designed hydraulic diameter of Dh = 508 \u00b5m. The material used for manufacturing the coupons was Inconel 718. The three different build directions studied are shown in Figure 2 (a-c) including horizontal, diagonal, and vertical. Each build orientation was defined by the position of the channel axes relative to the build plate. The channel axes were parallel to the build plate in the horizontal build direction, 45 degrees to the build plate in the diagonal build direction, and perpendicular to the build plate in the vertical build direction. Cylindrical shaped channels were chosen as the baseline to investigate build direction effects since it is a common design. In addition to the cylindrical channels, channels with diamond and teardrop shapes were also investigated",
            " Therefore, any feature of the test coupons that was less than 40\u00ba to horizontal required supports. Using this criteria, the upper surface of the horizontal cylindrical channels required supports; however, it would be difficult to remove these supports after the build was completed. As such, the horizontal surfaces of the channels were intentionally left unsupported to replicate situations where internal channels require supports, but cannot be accessed for support removal. Resulting support structures are shown in Figure 2. In addition to surfaces below 40\u00ba, supports were placed on other features, such as the coupon flanges required for the testing facility, to anchor them to the build plate to prevent distortion. For the study presented in this paper, the support structures were generated using a commercial STL editing and build preparation software. A state of the art DMLS machine was used to manufacture the coupons. Three duplicates of each test case were manufactured from the Inconel 718 powder for a total of 15 coupons"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure2.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure2.1-1.png",
        "caption": "Fig. 2.1 Conditional magnetic flux lines in the vicinity of the current-carrying induction machine windings (a) magnetic flux lines in the vicinity of the current-carrying stator winding; (b) magnetic flux lines in the vicinity of the current-carrying rotor winding; (c) magnetic flux lines in the vicinities of the current-carrying stator and rotor windings",
        "texts": [
            " The resulting fluxes \u03a61p and \u03a62p can be represented as the result of superposition of these \u201cpartial\u201d fluxes. We shall proceed under the assumption that these \u201cpartial\u201d fluxes are created by the current of each stator and rotor winding individually. The flux of the stator winding \u03a61p can be represented as the result of superposition of the total own flux\u03a61 (total flux of self-induction, Fig. 2.1\u0430) created by the current in the stator winding (in the absence of current in the rotor winding), and the flux of mutual induction \u03a621 (coupled with the stator winding, Fig. 2.1b) produced by the current in the rotor winding (in the absence of current in the stator winding), i.e., \u03a61p\u00bc\u03a61 +\u03a621. Analogously, the resulting flux of the rotor winding \u03a62p can be represented as \u03a62p\u00bc\u03a62 +\u03a612, where \u03a62 is the total own flux (total flux of self-induction) of the rotor winding (Fig. 2.1b) created by the current in the rotor winding (in the absence of current in the stator winding), and \u03a612 is the flux of mutual induction (coupled with the rotor winding, Fig. 2.1a) caused by the current of the stator winding (in the absence of current in the rotor winding). Thus the resulting fluxes of the stator and rotor windings are expressed as \u03a61\u0440 \u00bc \u03a61 \u00fe \u03a621 \u03a62\u0440 \u00bc \u03a62 \u00fe \u03a612 \u00f02:3\u00de 14 2 Fundamentals of the Field Decomposition Principle On the basis of equations (2.3), the resulting fluxes of the stator and rotor windings can be schematically represented in Fig. 2.1c. The fluxes\u03a61 and\u03a621 shown in the system of equations (2.3) induce the emfs E1 and E21 in the stator winding, where E1 is the total own emf (the total emf of selfinduction) and E21 is the emf of mutual induction of the stator winding. Therefore, the resulting emf of the stator winding is represented as E1\u0440 \u00bc E1 \u00fe E21 \u00f02:4\u00de The fluxes\u03a62 and\u03a612 used in the second equation of the system (2.3) induce emfs E2 and E12 in the rotor winding, where E2 is the total own emf (total emf of selfinduction) and E12 is the emf of mutual induction of the rotor winding",
            " Since the values of emfs E1 and E12 are the result of fluxes created by the currents in the stator winding, and as the values of emfsE2 andE21 are induced by the fluxes produced by the currents in the rotor winding, we can have E1\u00f0 \u00de \u00bc jx1I1; E21\u00f0 \u00de \u00bc jx21I2 E12\u00f0 \u00de \u00bc jx12I1; E2\u00f0 \u00de \u00bc jx2I2 \u00f02:6\u00de where x1 and x2 represent the total own reactance (total reactance of self-induction), and x21 and x12 also reflect the reactance ofmutual induction of the stator and rotorwindings. 2.1 An Induction Machine with a Single-Winding Rotor: Voltage Equations. . . 15 The emfs E1 and E2 represent the total fluxes\u03a61 and\u03a62 taking place in an electric machine under the unilateral power supply of its windings, and the emfs E12 and E21 reflect the fluxes of mutual induction\u03a612 and\u03a621, respectively (Fig. 2.1a, b). In this work, we consider a symmetrical electricmachinewith symmetrical power supply in its windings. Therefore, for ease consideration, the area of the magnetic field distribution can be limited to the single tooth division. The conditional magnetic field images corresponding to the equations (2.3) in Fig. 2.1 can then be represented in a more schematic and visual form in Fig. 2.2. In these figures, the stator and rotor teeth regions are subdivided into two areas: areas occupied by the windings and areas where slot wedges are located. In this case, the air gap is considered to comprise the two sub-layers with identical thicknesses equal to \u03b4/2. This provision is discussed below in greater detail. On the basis of expressions (2.4), (2.5), and (2.6), the resulting emfs of the stator and rotor windings (E1\u0440 and E2\u0440) take the form E1\u0440 \u00bc E1\u00f0 \u00de \u00fe E21\u00f0 \u00de \u00bc jx1I1 \u00fe jx21I2 E2\u0440 \u00bc E12\u00f0 \u00de \u00fe E2\u00f0 \u00de \u00bc jx12I1 \u00fe jx2I2 \u00f02:7\u00de The system of equations (2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000851_s12541-015-0332-6-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000851_s12541-015-0332-6-Figure2-1.png",
        "caption": "Fig. 2 Generated surfaces due to grinding wheel hub material using identical grinding parameters18 (Printed with permission from Elsevier)",
        "texts": [
            " In comparison with the steel substrate grinding wheel, the CFRP grinding wheel showed a smaller radial deformation. In addition, differences between the dynamic characteristics of these two grinding wheels also caused differences in their material removal mechanisms. Specifically, a crater-type structure was observed on the machining surface ground using the steel substrate grinding wheel, due to high-frequency vibration, whereas a groove-shaped structure was observed on the machining surface ground using the CFRP grinding wheel, due to the superior damping characteristics of CFRP (Fig. 2). However, the differing results were not effectively understood or analyzed. Additionally, the generation mechanisms and influences of using the CFRP grinding wheel were not discussed deeply enough. As a supplement, experiments were carried out to measure the amplitude and frequency at various points on or within the grinding wheel.50 Results showed that the amplitude at locations where the CFRP grinding wheel was fitted to the main shaft or came into contact with a workpiece was higher than the amplitude in the middle region of the grinding wheel substrate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000912_tfuzz.2015.2466111-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000912_tfuzz.2015.2466111-Figure2-1.png",
        "caption": "Fig. 2. Illustration of the Nyquist criterion (a) and the circle criterion (b)",
        "texts": [
            " We assume that: a) the transfer function G(s) has no common zeros and poles, b) the polynomial B(s) is stable, c) the polynomial A(s) does not have to be stable which means that it may have its roots in the origin or in the right half of the s-plane, d) the unstable system is stabilized by the linear feedback in the form of u(t) = re(t), where r > 0. To assess the stability of the system two frequency theorems are used \u2014 the Nyquist and the circle. These are based on the analysis of the Nyquist plots which are illustrated in Fig. 2. The Nyquist criterion [103] allows linear systems with delays to be studied. For these systems the analytical criteria of Routh or Hurwitz cannot be applied, as the transfer function of the plant is irrational. The Nyquist theorems for stable and unstable plants are recalled below. Theorem 1: The closed-loop system with the proportional controller u(t) = re(t) and a stable plant is stable if and only if the Nyquist plot G(j\u03c9) does not encircle the point (\u22121, j0) (Fig. 2a). The critical gain rc, i.e., the value of the gain r, for which the characteristics rG(j\u03c9) goes through the point (\u22121, j0) is determined by the formula rc = 1 |G(j\u03c9c)| , (2) where \u03c9c is the crossover frequency presented in Fig. 2a, such that the phase of G(j\u03c9) is equal to \u2212180\u25e6. Definition 1 (Hurwitz sector): The Hurwitz sector (0, rc) is the set of all the gains r \u2208 (0, rc) of the linear controller u(t) = re(t), for which the closed-loop system is stable. In case the plant is unstable, the Nyquist criterion requires knowledge of the number of P poles in the right half-plane. Theorem 2: The closed-loop system with the proportional controller u(t) = re(t) and an unstable plant is stable if and only if the Nyquist plot G(j\u03c9) encircles P times the point (\u22121, j0) counter-clockwise",
            " Below two forms of the circle criterion are given: for the sector [0, k], in which the circle is reduced to the line and for the sector [\u03b2, \u03b2+k] using the transformation (5). Theorem 3: A nonlinear system with the transfer function G(s) and nonlinearity bounded in the sector [0, k] is absolutely stable, if the following condition is met: inf \u03c9 Re [G(j\u03c9)] + 1 k > 0. (8) Geometrically, the circle theorem means that the Nyquist plot G(j\u03c9) lies to the right of the vertical line passing through the point \u22121/k (Fig. 2b). The maximum value of k may be determined from the formula k = \u22121 min \u03c9\u2208[0,\u221e) Re [G(j\u03c9)] . (9) If the nonlinearity f(e, t) satisfies the sector condition (4), then the circle criterion is used for the transformed transfer function (7) with the condition (6). Theorem 4: A nonlinear system with the transfer function G(s) and nonlinearity bounded in the sector [\u03b2, \u03b2 + k] is absolutely stable, if the following condition is met: inf \u03c9 Re [G1(j\u03c9)] + 1 k > 0, (10) where G1 is given by (7). In this case, the circle theorem means that the Nyquist plot G1(j\u03c9) lies to the right of the vertical line passing through the point \u22121/k"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000748_tie.2013.2276025-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000748_tie.2013.2276025-Figure11-1.png",
        "caption": "Fig. 11. Volumetric Joule losses pJ in billet for nnom = 1420 r/min.",
        "texts": [
            " Even for the above simplifying assumptions that reduced the time of computations more than ten times, the results exhibit a very good accordance with measurements, which is discussed in Section V-C. Both computations and measurements provided many results. In the next paragraphs, we will show and discuss the most important ones. Fig. 10 shows the distribution of magnetic flux density B in the system for the nominal revolutions of the billet nnom = 1420 r/min. Only very small areas near the corners of joints between the permanent magnets and the magnetic circuit are oversaturated. Analogously, Fig. 11 shows the distribution of the volumetric Joule losses pJ produced by the currents induced in the rotating billet at the same revolutions. In accordance with theory, its highest values are generated in the surface layers of the billet. Fig. 12 shows the distribution of temperature in the billet after 180 s of heating. It is obvious that the temperatures over its whole cross section do not differ more than by a few degrees of centigrade. This is caused by the very good thermal conductivity of aluminum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003765_iros.2018.8594316-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003765_iros.2018.8594316-Figure4-1.png",
        "caption": "Fig. 4: Detail of fingers.",
        "texts": [
            " Therefore, we made a novel thumb CM joint by the combination of three machined springs and realized the joint with DOFs including a joint with wide range of motion. And we applied a variable rigidity mechanism of agonistic tendon drive and placed one actuator to move this mechanism toward both of the fit of finger when grasping and the transmission of force with the side of fingers. The whole view and size of the hand we developed in this study are shown in Fig. 3. It is a five-fingered hand imitating human hands. It is connected to forearm through the wrist joint and it is moved by the muscles placed in forearm. The fingers of this hand is shown in Fig. 4. It has finger joints made of machined springs as with the hand of \u201cKengoro\u201d. The machined spring has flexibility and toughness because it is a kind of springs made of metal. In addition, it has an advantage that it can be connected firmly to other parts because it can be designed integrally with attachments. For these reason, it is suitable for finger joints of the hand, which is needed to support a large load and withstand impacts. Therefore, we made the fingers consists of three machined springs and the parts to connect springs imitating tendon sheath"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001536_s00170-018-2494-8-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001536_s00170-018-2494-8-Figure4-1.png",
        "caption": "Fig. 4 Illustration of a tool path used and b depth of cuts (DOC) during the side milling trials",
        "texts": [
            " Side milling operations were performed on each sample Table 3 GTAW process parameters for preparation of additively manufactured Ti-6Al-4V samples for side milling trials Sample ID Arc energy (J/mm) Specific deposition energy (kJ/g) Passes Total build height (mm) Wall_02 482 21.0 35 29.0 Wall_03 525 16.2 27 28.8 Wall_04 504 12.1 21 29.0 Wall_05 497 12.1 24 29.9 1 The results from this test trial are not presented as this test was carried out for experimental setup and calibration. using a 16-mm diameter SECO B39 JABRO HPM carbide end mill (designation JHP770160E2R250.0Z4A-SIRA) designed for machining of titanium alloys. The milling configuration, shown in Fig. 4, was selected to simulate the manufacture of deep pocketed, thin-walled features typical in structural aerospace components made from titanium alloys. Awater-oil emulsion was applied as a coolant to the cutting zone at high pressure (60 bar) using directed nozzles and also through the spindle of the cutting tool. Other parameters relating to the machining process are given in Table 4. The machinability of additively manufactured Ti-6Al-4V material produced by GTAW-wire based deposition was assessed relative to conventionally processed wrought material by comparing the machining forces during the side milling operation",
            " In the case of FX and FZ, a similar profile is observed with the magnitude of forces increasing to a maximum value whereas FY decreases to a minimum value until the tool becomes fully engaged with the workpiece. The forces then oscillate about a steady-state value for the remainder of the machining process. The initial peak in cutting forces is due to the tool path used, where the initial tool entry was through a circular arc originating at right angles to the wall and the tool exit path was parallel to the wall as illustrated in Fig. 4a. To achieve sequential full depth cuts, this tool path was offset by the required amount only in the radial (X) direction. It follows then that the maximum tool engagement is during the final section of the entry arc on the trailing side of the cutting tool. At this point, the radial depth of cut is equal to the full distance from the periphery of the tool to the outermost surface of the wall, producing the initial peak in cutting forces observed. L\u00f3pez de Lacalle et al. [27] also reported the initial peak in the cutting forces due to the tool engagement during an end milling process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002403_j.asoc.2015.10.068-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002403_j.asoc.2015.10.068-Figure12-1.png",
        "caption": "Fig. 12. Two-link robot.",
        "texts": [
            "in) j of the matrix are the same for all the subsystems, they can be all grouped as follows: J = \u2225\u2225\u2225\u2225\u2225 [ Y1 Y2 Ym p10 p20 pm0 ] \u2212 [ X ][ p1 p2 pm ]\u2225\u2225\u2225\u2225\u2225 2 J = \u2225\u2225Ya \u2212 Xap \u2225\u22252 nd the solution is still: p = (Xt aXa) \u22121 Xt aYa (36) A detailed proof is found in [1] . Illustrative examples In this section the proposed FLC-VSC is explained by two multivariable nonlinear examples of a two link robot and a thermal mixing ank. 178 B.M. Al-Hadithi et al. / Applied Soft Computing 39 (2016) 165\u2013187 E t m t s t w i 7 xample 7.1. Two link robot Let us consider the physical model of a two-link robot [37] (see Fig. 12), with each joint equipped with a motor for providing input orque, an encoder for measuring joint position and a tachometer for measuring joint velocity. The objective of the control design is to ake the joint positions 1 and 2 follow desired position histories d1(t) and d2(t), which are specified by the motion planning system of he robot. Such tracking control problems arise when a robot hand is required to move along a specified path, e.g. to draw circles. Let us uppose that 1 is the angle measured from the stable equilibrium position (inferior vertical) and 2 is the angle from the prolongation of he main link to the second one"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003182_j.compag.2017.08.016-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003182_j.compag.2017.08.016-Figure3-1.png",
        "caption": "Fig. 3. Illustration of the coordinates in the Serret-Frenet frame.",
        "texts": [
            " 2 presents the baits casting trajectory, which completely covers the given crab pond. The trajectory mainly determines two basic movements: straight line path following and spot turning. To save time, shorten the distance and reduce fuel consumption, the routes between turning points are approximated by straight rhumb lines (Li et al., 2009). Thus, straight line path following is the main objectives of the control system. The Serret-Frenet frame is widely used due to its flexibility in describing error dynamics (Li et al., 2009). Fig. 3 shows the Serret-Frenet frame used for path following of the operation boat. The origin of the frame fSFg is located at the closest point on the curve X from the origin of the body-fixed frame fBg. X is the given target path, e is defined as the distance between the origins of fSFg and fBg;wSF is the path tangential direction and w is the heading angle of the AOB. The dynamics of path-following errors in the Serret-Frenet frame is stated as follows: _e \u00bc u sin\u00f0 w\u00de \u00fe v cos\u00f0 w\u00de _ w \u00bc _w _wSF \u00bc j 1 ej \u00f0u sin\u00f0 w\u00de v cos\u00f0 w\u00de\u00de \u00fe r ( \u00f01\u00de where w \u00bc w wSF is the heading angle error, j \u00bc 0 is the curvature of the given path, u; v and r are the surge velocity, the sway velocity and the yaw angular velocity of the AOB, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001605_j.engfailanal.2018.09.032-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001605_j.engfailanal.2018.09.032-Figure1-1.png",
        "caption": "Fig. 1. a) Schematic of SLM equipment machine; and b) morphology of molten/solidified tracks.",
        "texts": [
            " The SLM samples were constructed using a laser machine operating through a Nd:YAG laser source with a 200-mm spot diameter, wavelength of 1.064 \u03bcm, and maximum output power of 100W. Galvanometric scanning mirrors were used to move the laser beam over the powder surface and draw every powder layer selectively. The powder deposition set-up includes a working basement in which a coater is used to store sequential powder layers along one direction. The working chamber uses a nitrogen atmosphere to prevent part oxidation and limit the initial oxygen level to 0.8%. Fig. 1a) provides a schematic of the SLM machine equipment used for the experiments. Several input parameters to be controlled and varied characterized the SLM process, and their scope was to reach an optimized parts quality. These include the laser (spot size, power, etcetera) and powder (shape, distribution, size, etcetera) parameters, hatch spacing (Hs), scanning speed (v), and layer thickness (LT). These parameters affect the molten/solidified track zone (Fig. 1b), thereby producing parts with different qualities. The layer thickness (LT) depends on the working basement stepping. In this study, it was set to 30 \u03bcm, which is the minimum step allowed by the employed SLM machine. Casalino et al. [18] used this step value to obtain effective adhesion between layers. All dog-bone specimens were manufactured following the ASTM E 8M [16] standard. The sizes (mm) and shapes of the samples are displayed in Fig. 2. Dog-bone specimens were deposited by the SLM system with three different processing parameter sets"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001779_j.proeng.2012.04.031-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001779_j.proeng.2012.04.031-Figure6-1.png",
        "caption": "Fig. 6. Aerodynamic torque applied on a shuttlecock, as it is not aligned with its velocity direction",
        "texts": [
            " The shuttlecock flip is possible because this object has distinguished centre of mass and center of pressure [6]. Actually, the shuttlecock cork is denser than its skirt so the center of mass is close to the cork for those objects. Meanwhile the aerodynamic center, where the drag is exerted, is close to the center of the volume that is to say close to the center of the skirt. When a shuttlecock is not aligned with its velocity direction the drag force, which is exerted on the aerodynamic center, submits a stabilizing torque to the shuttlecock. This stabilizing torque, reported on figure 6, explains why the shuttlecock flies the cork ahead and flips after racket impact. Writing torque equilibrium on the shuttlecock provides a prediction for reversing, oscillating and stabilizing time. For a high clear, the time of an exchange is typically 2 s. In this case the stabilizing time (about 0,2 s) is short compared with this time of exchange. We can neglect the flip in the dynamic of the shuttlecock and the approach of the second section, which assumes a constant cross section, is validated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001418_s12555-017-0081-7-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001418_s12555-017-0081-7-Figure1-1.png",
        "caption": "Fig. 1. Quadrotor UAV.",
        "texts": [
            " Guidelines for choosing controller parameters of the proposed control system are as follows: 1) The value of k j,i j in the sliding surface denotes the nonlinear slope of the super-twisting sliding surface which determines the speed of reaching mode into the nonlinear super twisting surface from any initial point. 2) The value of 1/2\u2264 \u03b3 j,i j < 1 determines the finite time convergence to steady-state, where smaller \u03b3 j,i j increases the convergence speed. If \u03b3 j,i j = 1, the super-twisting sliding surface becomes the normal first order sliding surface. 3) The value of \u03b2 j,i j has a similar role to proportional gain in a PID controller. 5. APPLICATION EXAMPLE Fig. 1 shows a quadrotor UAV as an application example for the proposed control scheme. We designed four control systems to compare control performances BSC, STA-BSC, DSC , and STA-FBSC systems. Model of the translational and attitude dynamics in the navigation frame were described as follows: x\u0308 = 1 m (cos\u03d5 sin\u03b8 cos\u03c8 + sin\u03d5 sin\u03c8)uz \u2212 K1x\u0307 m , (70) y\u0308 = 1 m (cos\u03d5 sin\u03b8 sin\u03c8 \u2212 sin\u03d5 cos\u03c8)uz \u2212 K2y\u0307 m , (71) z\u0308 = 1 m (cos\u03d5 cos\u03b8)uz \u2212g\u2212 K3z\u0307 m , (72) \u03d5\u0308 = \u03b8\u0307 \u03c8\u0307 Jy \u2212 Jz Jx + Jr Jx \u03b8\u0307\u2126r + l Jx \u03c4\u03d5 \u2212 K4l Jx \u03d5\u0307 , (73) \u03b8\u0308 = \u03c8\u0307\u03d5\u0307 Jz \u2212 Jx Jy \u2212 Jr Jy \u03d5\u0307\u2126r + l Jy \u03c4\u03b8 \u2212 K5l Jy \u03b8\u0307 , (74) \u03c8\u0308 = \u03d5\u0307 \u03b8\u0307 Jx \u2212 Jy Jz + l Jz \u03c4\u03c8 \u2212 K6l Jz \u03c8\u0307, (75) where m = 2 kg is the quadrotor mass, Jx = Jy = 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003611_j.mechmachtheory.2018.09.009-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003611_j.mechmachtheory.2018.09.009-Figure3-1.png",
        "caption": "Fig. 3. FE contact model for the cam-tappet.",
        "texts": [
            " The bending deformation is also a major component during the rise and return phases. With the development of the FE method, contact stiffness between complex parts can be obtained by quasi-static contact analysis given a high precision and general applicability [38] . Therefore, the FE contact analysis is conducted to calculate the dry contact stiffness in this work. The contact, bending and shear deformations are all considered in this model. The FE quasi-static contact model for the cam-tappet is demonstrated in Fig. 3 . Directions y 1 and z 1 are the axial directions of the camshaft and tappet, respectively. Direction x 1 refers to the radial direction of the camshaft. Torus 1 and 2 are the contact surfaces between the camshaft and the bearings and are applied to define constraints. In Eq. (5) , the K d comprises the cam stiffness K cam ( \u0303 t) and tappet stiffness K tappet ( \u0303 t) . Six DOFs of nodes on Torus 1 and 2 are fixed for the constraints. The other DOFs of nodes on the top surface, except the translation freedom along z 1 , are also restricted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000850_s0081543815080210-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000850_s0081543815080210-Figure1-1.png",
        "caption": "Fig. 1. The forces acting on the cart and the pendulum.",
        "texts": [
            "1) approach a target on a fixed time interval and the optimal time problem for system (1.1) in the case when it is stationary. 3. SIMULATION OF APPROACH PROBLEMS FOR MECHANICAL CONTROL SYSTEMS In this section, we consider some control systems on a finite time interval [0, T ] and the problems of approach to a point target at time T for these systems. To solve these problems, we apply the schemes and algorithms described in the preceding sections. Example 1. Consider an inverted pendulum with the point of suspension on a cart moving along a horizontal plane (see Fig. 1). The cart has variable mass M(t), t \u2208 [0, T ], that decreases according to a given law as time t increases. The cart is acted upon by the force of gravity # \u2014 M(t)g applied to the center of gravity of PROCEEDINGS OF THE STEKLOV INSTITUTE OF MATHEMATICS Vol. 291 2015 the cart, the friction force with constant friction coefficient k > 0, the drag force #\u2014 F (t), t \u2208 [0, T ], the reaction force due to the expulsion of a part of the mass of the cart (see, for example, [9]), as well as the reaction force from the supporting horizontal plane and the pendulum"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002445_ijmmme.2016010103-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002445_ijmmme.2016010103-Figure6-1.png",
        "caption": "Figure 6. Temperature profile variation on gearbox surface- Isothermal gear oil bath temperature (1000C)",
        "texts": [
            " The lay/counter shaft also shows variation of temperature (346.07-352.38) k for h400w/m 2 k. Figure 5 (e, f) shows very similar temperature profile for gear train. The 3rd gear profile shows temperature increment in red hues and this area is prone for thermal stresses more for (h500w/m 2 k, h600w/m 2 k) convective heat transfer coefficient values. In figure 7, h400 and h500 shows the gear train temperature profile variation. For second part of study the gear oil temperature is increased to 1000 C. Figure 6 shows the temperature profile of gearbox surface at gear oil temperature of 1000C. Figure 6(a) shows the temperature profile of gear train assembly at h100w/m 2 k convective heat transfer coefficient (h) value. The minimum temperature is 342.81 k and maximum is 364.02 k. Maximum temperature effect is found on same place as for h100w/m 2 k at 800C on counter shaft right end in red hues. Red hues shows hot areas subjected to thermal stresses and deformation. Figure 8, h100 shows the gradual increment in temperature profile of gear train at different points. Figure 6(b) explains the temperature variation in transmission gear train at h200 w/m 2 k. As the value of h increases there is increase in temperature of gear train by 4.76 k, mention by red hues. The high hues temperature region is found at fixed portion of right side end of counter shaft. Figure 6(c) shows dark yellow hues on 3rd gear. The dark yellow hues designate increment in temperature. The maximum value of temperature reached to 370.09 k. The right side fixed counter shaft end temperature is maximum for h300w/m 2 k. The temperature profile varies between (342.8-370.09) k. In figure 8, h300 (3) shows the temperature linear temperature variation at different point on gear train. Figure 6 (d, e, f) shows gradual change in temperature. Gear profile of 2nd and 3rd shows red hues of hot areas. In figure 8, h600 shows the gear train temperature profile variation. Multi speed transmission gearbox and gear oil analysis is highly nonlinear problem. From FEA analysis (Figure 5 and 7), when h is 100 w/m2k and gear oil temperature is constant at 800C, the gear train profile temperature varies very gradually and linearly. As we increase the value of convective heat transfer coefficient to 200 W/m2 k the difference in maximum temperature is only 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003728_tsmc.2018.2884619-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003728_tsmc.2018.2884619-Figure5-1.png",
        "caption": "Fig. 5. Discrete state space is built upon exteroceptive information in a stable swinging movement. The SOM training set is composed of the torso angle with the vertical axis and its derivative in time, it is represented by black dots in (a). SOM neural network (blue nodes and fixed connections shown in red) represents the state space of a swinging movement after learning in (b). Gray lines are the interpolation of the training set over time.",
        "texts": [
            " We take advantage of this to modulate the stream of exteroceptive information as a dynamic state space. This state space is a discrete representation of the continuous state of the robot, it preserves limit cycle characteristics of a stable rhythmic movement. A SOM network has been developed so that its neighborhood function arranges all neurons in a circular topology. The input vector is composed of the vertical angle of the robot\u2019s torso \u03b8 , and its rate of change \u03b8\u0307 during a stable swinging movement; see Fig. 5(a). With predefined N-neurons, the output of the network presents a state space of N distinct state neurons. Activation of state neurons during a stable rhythmic movement follows a neighboring order representing the dynamic change of the robot\u2019s state. Each state will be activated for a specific time before the activation moves to the next state. Accordingly, the duration of activation at each state reflects the dynamics of the robot while swinging with a fixed frequency. If the swinging frequency is disturbed, SOM neurons may keep activating in the same sequence for certain time but the duration of activation will be effected at the time of disturbance. This duration is used as a criterion for stable swinging. Fig. 5(b) shows a state space resulting from the training of a circular SOM network composed of 8-neurons. The duration of activation differs from one state to another state during a stable swinging movement. For instance, we obtained a minimum duration of four cycles (Tmin = 160 ms) and a maximum duration of ten cycles (Tmax = 400 ms) for 2.5 Hz swinging frequency. The activation sequence of state neurons in a stable rhythmic movement stays strictly on the limit cycle as explained in Section III-C. An external perturbation would take the robot motion away from the stable rhythmic movement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001649_j.sna.2018.10.031-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001649_j.sna.2018.10.031-Figure8-1.png",
        "caption": "Fig. 8. Experiment",
        "texts": [
            " (21) he open-loop transfer function of the system in Fig. 3 is given by q. (22) open = a(s) [ Cf (s) + Cp(s) ] e\u2212T0s s2 1 \u2212 Cf (s) a(s) e\u2212T1s s2 . (22) sed on acceleration open-loop fusion. The closed-loop system with feed-forward control should have a phase margin larger than 45deg of the open-loop transfer function, we can get the following: arg[Sopen|s=jwc ] + 180\u25e6 \u2265 45\u25e6. (23) Consequently, while implementing the Kalman filter, it is also necessary to meet the requirements of Eq. (23) 5. Experimental verification Fig. 8 shows the experimental devices. The CCD-based fine tracking system is a two-axis system. Due to the symmetry of the two axes, this paper focuses on only one axis. The light source emits light to simulate the target. The motion of the target is simulated by rotating the target mirror. The MEMS accelerometers mounted on the back of the tracking mirror measure the platform acceleration and achieve the inner acceleration loop. The CCD detects the light-of-sight(LOS) error of the simulated target and implements the outer position loop"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002668_1350650115619610-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002668_1350650115619610-Figure3-1.png",
        "caption": "Figure 3. Finite element mesh of thrust bearing with circular shape recess.",
        "texts": [],
        "surrounding_texts": [
            "and\n@ p @ _h \u00bc\n@p1 @ _h @p2 @ _h\n..\n.\n@pi @ _h\n..\n.\n@pn @ _h\n2 666666666664 3 777777777775\nSolution procedure\nThe elastohydrodyanmics problem is a problem of fluid\u2013structure interaction. For the coupled solution of governing equations, the adopted approach is shown in Figure 2. The problem requires the simultaneous solution of the Reynolds and elasticity equation. The solution procedure of this method involves Newton\u2013Raphson method and other iterative method. In this bearing deformation is computed to get fluid-film thickness variation on bearing surface. To compute static pressure and fluid-film thickness, fluid-film pressure and bearing pad deformation are initialized zero. The iterations are continued until the following convergence criterion is not satisfied\ntol \u00bc 0:0014 max p0 k\np0\nk\n! \u00f031\u00de\nThe complete solution procedure is applied in the following steps:\n1. Discretize the two-dimensional lubricant flow field into 4-noded quadrilateral isoparametric elements and discretize bearing pad geometry into 8-noded trilinear brick element. 2. Initialize the value of fluid-film pressure, fluid-film pressure gradient, and bearing pad deformation on nodes. 3. Identify the type of lubricant i.e. Newtonian or non-Newtonian and also check the type of the bearing pad is rigid or flexible.\n4. Compute the deformation of the thrust bearing pad by using equilibrium conditions by using equations (21) and (22). 5. Compute fluid-film thickness and compute nonNewtonian coefficient F0, F1, and F2. 6. Apply boundary conditions to global fluidity matrix. 7. Global system of equation for fluidity matrix is solved by using Gauss\u2013siedel iterative method and global stiffness matrix for bearing pad deformation is solved by using the backward substitution method. 8. Repeat steps 3 to 7 until the convergence criteria are not satisfied. 9. Once the convergence criteria is satisfied. Compute the static and dynamic performance characteristic parameter.\nResults and discussion\nOn the basis of the above solution methodology of nonlinear iterative finite element method a MATLAB program has been developed. The present simulation involves fluid\u2013structure interaction. The average computation time for a single solution is about 6 hrs. As there is no finite element formulation for elastohydrostatic lubrication with non-Newtonian\nTable 1. Bearing performance parameter.\nCs2\u00bc 30 Cs2\u00bc 40\nPerformance characteristics Ref.1 Present Variation (%) Ref.1 Present Variation (%)\npoc 0.9298 0.9297 0.011 0.9520 0.9520 0.00 F0 1.5803 1.5808 0.0316 1.6181 1.6187 0.037 S 0.3330 0.3332 0.0601 0.2327 0.2329 0.0859\nC 0.7365 0.7351 0.1901 0.6907 0.6892 0.2172\nat UNIV CALIFORNIA SAN DIEGO on December 12, 2015pij.sagepub.comDownloaded from",
            "lubricant, it has not been reported in literature (to best of the author\u2019s knowledge). Therefore, to check the validity of the developed algorithm and methodology a check has been performed with the results presented\nin the open literature. Therefore, the results are compared with the analytical solution of the hydrostatic thrust bearing operating with Newtonian lubricant ( \u00bc 0). As shown in Table 1, results have good\nat UNIV CALIFORNIA SAN DIEGO on December 12, 2015pij.sagepub.comDownloaded from",
            "agreement with the previously published analytical results and maximum percentage variation in the results is 0.22%. The results also compared with tilted pad hydrostatic thrust pad bearing and presented in Table 2. The maximum variation found in the lubricant flow rate is 4.11%\nThe values of bearing operating and geometric parameter are judiciously selected on the basis of available published literature12,34,42 and are presented in Table 3. As shown in Figures 3 to 6, the discretization of lubricant flow field and discretization of thrust pad bearing domain has been performed by using twodimensional quadrilateral and three-dimensional 8- noded isoparametric brick elements, respectively. All recess shape has equal value of Ab\nAoc \u00bc 4. To discretize\nthe flow field domain of circular, rectangular, square, and elliptical shape recess, meshes of 1150, 1128, 1069, and 1255 elements have been used in the analysis. To discretize the deformation field of circular, rectangular, square and elliptical shape recess, meshes of 11500, 11280, 10690, and 12550 elements have been used in the analysis. These meshes are developed with\nthe help of ANSYS 12.0 and imported in MATLAB 2012a software.\nTo visualize the physical effect of the elasticity of the bearing pad a fluid-film pressure distribution is presented in Figures 7 to 10. The fluid-film pressures are presented for the selected parameters. The thrust bearing operated with Newtonian lubricant has the maximum value of fluid-film pressure and the bearing operating with non-Newtonian lubricant with rigid thrust pad has the minimum value of fluid-film pressure. The pocket pressure of rectangular shape recess is the minimum and pocket pressure of circular shape recess is the maximum. Fluid-film pressure distribution is significantly affected on the consideration of the nonNewtonian and elastohydrostatic. In Figure 11, the results are presented for the all types of recess shape. The thrust pad having circular shape recess has maximum deformation and thrust bearing having rectangular shape recess has minimum deformation. In Figures 12 to 16, results are presented for the Newtonian and psuedoplastic lubricant for different values of tilt parameter and restrictor design parameter.\nat UNIV CALIFORNIA SAN DIEGO on December 12, 2015pij.sagepub.comDownloaded from"
        ]
    },
    {
        "image_filename": "designv11_13_0000637_1350650114538779-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000637_1350650114538779-Figure1-1.png",
        "caption": "Figure 1. Geometry of dimple textured slider bearing.",
        "texts": [
            " In the present study, an attempt has also been made to study the effect the direction of texturing (forward and backward) on the dimple textured slider bearing in addition to effect of inclination. A theoretical model has been developed to study the slider bearing property, for a textured surface consisting of spherical micro-dimples. Here, the slider bearing consists of two parallel plates. The clearance between two plates is h0. The upper plate is moving with a velocity U, when the lower plate is fixed as shown in Figure 1. The lower plate consists of spherical micro-dimples as shown in Figure 2(a). The dimple base radius is rp and the maximum dimple depth is hp. At any distance the height of the fluid film (local film thickness) is h (x,z). Dimensionless clearance, , dimensionless dimple depth, hp, dimensionless length x, dimensionless local length, x , dimensionless local width, z , dimensionless bearing length, L, dimensionless bearing width, B, dimensionless local height, h, and textured portion of slider width, , are given as \u00bc h0 2rp , hp \u00bc hp h0 , x \u00bc x rp , x \u00bc x rp , z \u00bc z rp , L \u00bc L rp , B \u00bc B rp , h \u00bc h h0 , \u00bc Bp B The spherical dimple textured surface is developed in the periodic repetition of a small dimple in a small imaginary square cell of arm length of 2r1 on the textured surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003422_s12206-018-0307-5-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003422_s12206-018-0307-5-Figure6-1.png",
        "caption": "Fig. 6. The compression test of: (a) Solid tire; (b) pneumatic tire on the rotating drum.",
        "texts": [
            " The both types of tractor tire were performed the drum testing under the same conditions. The drum testing was set up the rolling tire speed at 10, 20, 30, 40 and 50 km/h by supporting a constant compression load of 400 kg. The compression loads were also variable by setting up at 200, 300, 400, 500, 600 and 700 kg under the constant rolling speed of the tires at 20 km/h. The compression load had been restricted in the horizontal direction by pressing tires on the rolling drum using the tiremounting arm. Fig. 6 shows the drum testing of tires using the instrumented drum wheel. The force sensor at the mounting arm was used to measure the normal contact force between drum and tire. The error was less than 0.5 N for the contact force measurement. The tires would be rotated to warm on the drum about 30 minutes before the drum testing at each compression load and speed. The testing time was at 900 sec for the certain condition of the steady state. The measuring time was started after the testing time of 300 sec; therefore, the recording time was 600 sec for each drum testing condition which ensured the rolling tires responding in the steady state range of time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003887_j.ymssp.2019.04.039-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003887_j.ymssp.2019.04.039-Figure5-1.png",
        "caption": "Fig. 5. (a) FEM model, (b) scheme of excitation points and response point.",
        "texts": [
            "58) and (106.2, 108.2) features the order pair, which comes from the convolution of Stage 1 meshing components located at 53.6 order and 107.2 order and the components pair of the projection function located at order(-1, rameter. 1). The above confirm the analytical results in section 2.2. Additionally, the amplitudes decrease with the increase of frequency, as shown in Fig. 4(f). In order to obtain the FRF of the transfer path, a 3-D model is built with the finite element method (FEM) as shown in Fig. 5(a). The model establishes with 278,539 nodes and 59,587 tetrahedral elements to include Stage 2 ring gear and the gear casing parts. Modal analysis is conducted to obtain the one-dimensional FRF cluster by imposing sweeping force individually on the ring gear and sampling the vertical response of the Stage 2 ring gear at the transducer location u \u00bc p=2. The input locations are assigned with the 28 circumferentially distributed on the meshing points of the ring gear. The two- dimensional FRF is obtained by composing the one-dimensional FRF cluster with the time-angle map reference of the carrier"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003283_0954406217745336-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003283_0954406217745336-Figure6-1.png",
        "caption": "Figure 6. The force model of rail.",
        "texts": [
            " The rail\u2019s elastic deformation was mainly provided by rail pads. The system\u2019s mechanics model is shown in Figure 2, where kp is the stiffness of the rail pads; cp is the damping of the rail pad, which is acquired by experiments. Rail systems are very complicated. Actual numerical computation always involves a simplified supported beam model with finite length. Practical results have shown that such processes can completely satisfy the precision demand of engineering. This paper uses the Euler beam,24\u201326 and the force model is shown in Figure 6, where P is the rail\u2019s contact force, the right wheel\u2019s is N0Ry. The wheels\u2019 speed is v; FRi(i\u00bc 1,. . .,N) is the pivot i\u2019s reaction under rail, N is the total number of pivots under the rail within the total length of that rail; x0 is the initial distance from the wheel to the left fixed end; xi is the section i\u2019s coordinate, the original coordinate O was on the left side of the simply supported end. The rail\u2019s vertical vibration displacement is Yr(x,t). According to the basic characteristic of Euler beam unit, the vibration differential equation is as follows EI @4Yr\u00f0x, t\u00de @x4 \u00femr @2Yr\u00f0x, t\u00de @t2 \u00bc XN i\u00bc1 FRi\u00f0t\u00de \u00f0x xi\u00de \u00fe P \u00f0x xw\u00de \u00f027\u00de where mr is the mass of the rail with unit length; E is the rail\u2019s elasticity modulus; I is the rail\u2019s cross section inertia; (x) is the unit pulse function, when x\u00bc 0, (x)\u00bc 1, and (x)\u00bc 0 under other circumstances"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure1-1.png",
        "caption": "Fig. 1. Traditional undercutting \u2014 type I: a) undercutting; b) non-undercutting; and c) boundary case.",
        "texts": [
            " That is why in the present paper a new generalized approach for defining undercutting of teeth is proposed, where simultaneously with the traditional boundary case (undercutting \u2014 type I) twomore boundary cases, defined as \u201cundercutting\u2014 type IIa\u201d and \u201cundercutting\u2014 type IIb\u201d, are taken into consideration. In these two additional cases the undercutting is done by the rack-cutter fillet but not by its rectilinear profile. Hence the undercutting type IIa is characterized by the decrease of the tooth thickness in their dedendum without cutting the involute profile while in the presence of undercutting type IIb a part of the initial section of the involute profile is cut. The undercutting of teeth\u2014 type I of the involute gear is obtained when at its meshing with the rack-cutter (Fig. 1a), points of the rectilinear profile KE of the tooth cutter are situated under the line of action AB [1,2]. Then the trajectories of these points (lines parallel to the rack-cutter centrode n-n) do not cross the line of action AB (they cross its extension), and the rectilinear segment \u041a\u0415 of the rack-cutter becomes a non-operating profile that penetrates into the bottom of the processed tooth and cuts a part of an involute profile. In order to avoid the undercutting \u2014 type I, it is necessary that the tip-line g-g of rectilinear profile of the rack-cutter crosses the line of action AB (Fig. 1b). Otherwise, when the crossing point L lies outside the line of action on the extension of the line PA (Fig. 1\u0430), the cut teeth are undercut. This means that the condition for non-undercutting \u2014 type I generally can be expressed by the inequality PA\u2265PL; \u00f01\u00de king into account the specified distances marked on Fig. 1 PA \u00bc r sin\u03b1 ; PL \u00bc ha\u2212X\u00f0 \u00de= sin\u03b1 , the traditional condition for and ta non-undercutting \u2014 type I finally is written in the following way [2,18,19](2) \u00f02\u00de In the inequality (2) x=X/m is an addendummodification coefficient of the rack-cutter, ha\u204e=ha/m is an addendum coefficient of basic rack tooth, z represents the number of teeth, \u03b1 is the profile angle of the rack-cutter, andm is themodule of the gear (m=2r/z). Minimum addendummodification Xmin of the rack-cutter, eliminating the undercutting \u2014 type I, is defined by the equation Xmin \u00bc xminm \u00bc h a\u22120:5z sin2\u03b1 m : \u00f03\u00de Xmin corresponds to the so-called boundary case \u2014 type I (Fig. 1c), where the tip-line g-g passes through the boundary point A (the point where the line of action contacts the base circle of a radius rb). As already mentioned, the undercutting \u2014 type II is caused by the rack-cutter fillet AF (Fig. 2) in the process of tooth cutting. When this fillet is a circle of a small radius \u03c11 (Fig. 2a) the teeth cut are not undercut. Then the gear fillet fb does not cross the radial line Ob, passing from the center O to the starting point b of the involute profile ba (at X=Xmin point b lies on the base circle of a radius rb)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000017_icsmc.2012.6378165-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000017_icsmc.2012.6378165-Figure4-1.png",
        "caption": "Fig. 4. Joint model from motor circuit to joint link",
        "texts": [
            " 3 will start falling down freely by an external force. If there is no friction (damping) at the joint, ZMP will stay at the ankle joint as shown in Fig. 3(a). On the other hand, if the friction of the joint is very high, then the joint will be locked and ZMP will be located on the edge of the foot as in Fig. 3(b). Fig. 3(c) shows that ZMP changes depending on the joint friction. The joint model is briefly described in order to derive a friction model. Considering only one joint, the joint model from motor circuit to joint torque in Fig. 4 results in the following equations\u23a7\u23aa\u23a8 \u23aa\u23a9 V = Ri + Ldi dt + Kbq\u0307m \u03c4m = Jmq\u0308m + n\u03c4l + \u03c4f Jlq\u0308l = \u03c4l. (11) In Fig. 4 and (11), V , i, R, L, Kb are voltage, current, resistance, inductance, back-emf constant and torque constant of the motor circuit, respectively. Also qm, ql, Jm, Jl, \u03c4m and \u03c4l denote motor angle, joint angle, motor inertia, link inertia, motor torque and joint torque, respectively. To derive a friction model, it is necessary to convert (11) into an equation with the variables that can be measured or controlled. We assume the following. 1) The inductance of motor circuit is very small. (R,Kt L \u21d2 L \u2248 0) 2) The motor torque is proportional to motor current multi- plied by a torque constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000328_j.ibiod.2014.05.021-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000328_j.ibiod.2014.05.021-Figure1-1.png",
        "caption": "Fig. 1. A schematic of the single-ch",
        "texts": [
            " Titanium wire was fitted to the body of the MFC and the end plate with a hole for air contact. The wire was then sealed with a conductive silver epoxy for connection to a fixed external circuit resistance at 1000 U. A thermometer was embedded in the reactor to monitor the temperature of the anode chamber. Silicon gaskets were placed between the end plates and the reactor body as well as the electrode and membrane in order to prevent water leakage. All parts were fabricated with rods and nuts as shown in Fig. 1. Return activated sludge was obtained from the Jungnang sewage treatment plant in Seoul, Republic of Korea as the inoculum. The sludgewas rinsedwith a phosphate buffer solution three times. The pretreated sludge and a medium solution containing 0.5 g L 1 acetate were injected into the MFC at a ratio of 1:1 for the start-up stage. The chemicals used for the phosphate bufferedmediumwere as follows (grams per liter of deionized water, g L 1): NaH2PO4$H2O, 2.69; Na2HPO4, 4.33; NH4Cl, 0.31; KCl, 0.13, as reported by Kim et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000444_0954406214524743-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000444_0954406214524743-Figure2-1.png",
        "caption": "Figure 2. The geometry of the big-end bearing.",
        "texts": [
            " The final Lagrange equation can be written as: mp \u00femb b1mb sin b1mb sin Ib \u00fe b21mb \" # \u20aczp \u20ac \" # \u00fe 0 _ b1mb cos 0 0 \" # _zp _ \" # \u00bc Fzb Fg Fybb cos Fzbb sin \" # \u00f015\u00de at MICHIGAN STATE UNIV LIBRARIES on February 28, 2015pic.sagepub.comDownloaded from where Fg is the combustion force acting on the top of the piston Hydrodynamic force in the big-end bearing Reynolds equation is a classic equation to solve the hydrodynamic forces in the lubrication oil. Pinkus and Sternlicht14 have made a very detailed analysis of the hydrodynamic forces of different journal bearings. The geometry of the big-end bearing is shown in Figure 2. There are two coordinate systems. One is the global YZ system and the other is related to the eccentricity line rt system, based on the diameter through the point of minimum oil thickness. The X direction is inline with the length of the bearing. is the coordinate angle between the positive r direction and the positive Y direction. Because the load on the big-end journal bearing is a variable rotating load, should be a combination of attitude angle and rotation angle of the load. e is the eccentricity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000554_cdc.2012.6426837-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000554_cdc.2012.6426837-Figure2-1.png",
        "caption": "Fig. 2. Quadrotor configuration [7]",
        "texts": [
            " rank ([B(ta) |Bs]) = rank   2 7 3 0 1 0 0 1 0 1 0 0 0 0 1 0  = 4 (30) Therefore, it can be concluded that the obtained slack variable matrix Bs satisfies the purpose of the slack variable introduction. To verify the performance of the proposed controller, the landing scenario of a quadrotor UAV system is considered. The quadrotor UAV system is one of the control affine underactuated nonlinear systems. The quadrotor UAV system has four rotors, and each rotor produces a lift force and moment. Two diagonal rotors (1,3) rotate in the same direction, whereas the others (2,4) operate in the opposite direction, as shown in Fig. 2. Varying the pair of the rotors\u2019 speed produces roll, pitch, and yaw motion. Due to the geometric symmetric property of the quadrotor UAV system, the coupling inertia moments are assumed to be zero. Then, the equations of motion can be represented as x\u0308 = u1 (cos\u03d5 sin\u03b8 cos\u03c8 + sin\u03d5 sin\u03c8) (31) y\u0308 = u1 (cos\u03d5 sin\u03b8 sin\u03c8 \u2212 sin\u03d5 cos\u03c8) (32) z\u0308 = u1 (cos\u03d5 cos\u03b8)\u2212g (33) \u03d5\u0308 = u2l (34) \u03b8\u0308 = u3l (35) \u03c8\u0308 = u4 (36) where \u03d5 , \u03b8 , and \u03c8 are roll angle, pitch angle, and yaw angle, respectively, g denotes the gravitational acceleration, and l represents the length from the center of the quadrotor to the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001808_1.4037667-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001808_1.4037667-Figure6-1.png",
        "caption": "Fig. 6 Three-pad HAFB with three orifice tubes",
        "texts": [
            " Figure 4 shows the top foil with welded orifice tube. When the top foils are assembled on to the bearing sleeve, the orifice tubes are located at h \u00bc 60deg, h \u00bc 180deg (direction of gravitational loading), and h \u00bc 300deg. Figure 5 shows orifice tubes configuration for the three-pad HAFB. Orifice tubes are connected to the main supply pressure line, and the controlled hydrostatic injection is achieved by controlling the air pressure to the orifice tube located at h \u00bc 180deg through a separate on/off valve. Figure 6 shows the three-pad HAFB. The rotor is constructed with several components. Bearing journals are at each side of induction motor element, and thrust runner and end cap (with the same weight as the thrust runner) are at the each ends and they are all assembled with two tension bolts, ensuring the first bending critical speed of the rotor assembly far above the maximum operating speed. Table 1 summarizes the overall geometries and parameters of the HAFBs and the rotor. Figure 7 depicts the cross section view of the test rig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002131_idam.2014.6912676-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002131_idam.2014.6912676-Figure1-1.png",
        "caption": "Figure 1 \u2013 The Knowledge scope of mechatronic system",
        "texts": [
            "EYWORDS: Mechatronics, Multicriteria, Multi-disciplinary, Concurrent Design, Integration, Optimization. Mechatronic systems are seen as a combination of cooperative mechanical, electronics and software components aided by various control strategies (Figure 1). They have led engineering design into a new era by integrating the most advanced technologies with the best design schemes. Design of a wide variety of products such as transportation systems, aircrafts, construction machines, robots or even home appliances are now considered within the area of mechatronic systems design. The term Mechatronics was introduced in late 1960s, but the significant growth of the field was observed in the 1990s (Behbahani 2007). Mechatronic systems are often highly complex, because of the high number of their components, their multi-physical aspect and the couplings between the different engineering domains involved which make the design task very tedious and complex"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002568_jahs.59.042008-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002568_jahs.59.042008-Figure3-1.png",
        "caption": "Fig. 3. Control mode architecture.",
        "texts": [
            " Furthermore, it was preferred to use the piloted handling control laws for maneuvering and flight away from hover because the motions of the load are not critical in this regime, as long as they are stable. It was also apparent that the load damping control laws should be used in hold modes when the pilot is not in the loop. In addition, load damping should preferably be applied at very low speed and hover, where the load motion is most important because the load is most likely to be placed on the ground from this flight condition. The moding architecture that was developed to implement the tasktailored control strategy is shown in Fig. 3. The system features attitude command, velocity hold, automatic deceleration, and position hold (PH) modes. The load damping (cable angle/rate feedback) control laws come on during the deceleration and PH modes, as shown in Fig. 3(b), which occur at low speed when the pilot has the cyclic in the center detent and therefore is not actively maneuvering the aircraft. The pilot handling (cable angle feedback) control laws are active during maneuvering when the stick is out of detent and also in the velocity hold mode. The maximum speed for the control law is 40 kt because this study focuses only on hover/low-speed handling qualities and load placement tasks. Descriptions of all the modes are given in Table 1. This control law scheme for CAF allows the pilot to maneuver with good handling qualities to the load set-down location",
            " If at any time the pilot wants to maneuver the helicopter, the control system automatically switches to the handling qualities cable angle/rate gains when the stick leaves the detent position. The load damping returns when PH or automatic deceleration is reenabled. Thus, the task-tailored control law combines both the load damping and pilot handling control laws into a multimode control law architecture and will be henceforth be referred to as the CAF control law. An optimized baseline system with fuselage feedback only was developed with the same command and hold modes but no load specific modes, as shown in Fig. 3(a) and described in Table 1. The baseline system does not switch gains in a task-tailored way because there are no cable feedbacks. This system provides a well-designed and fair basis for comparison with the CAF system, since it has the same architecture (without cable feedback) and was designed against the same specifications. This control law extends the baseline attitude command/attitude hold system described in the preceding trade-offs section of this paper to include velocity and PH modes and optimizes the fuselage gains for this configuration. This control law is referred to as the optimized baseline control law (OBL). The task-tailored control laws with modes shown in Fig. 3 were implemented with architecture shown in Fig. 4 for the lateral and longitudinal axes. The basic attitude command control system is an explicit model following control law (Ref. 17). It is used when the stick is out of the detent position. The velocity and PH modes were implemented with nested velocity and position feedbacks as shown in Fig. 4. The load feedback is inertial (\u03c6c,\u03b8c) as opposed to relative to the aircraft. As an example, the lateral inertial cable angle \u03c6c and relative cable angle \u03c6c are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001197_j.ins.2017.08.090-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001197_j.ins.2017.08.090-Figure6-1.png",
        "caption": "Fig. 6. State trajectory of the first component of x\u0304 h (k, l) in Example 2 when w (k, l) \u2261 0 .",
        "texts": [],
        "surrounding_texts": [
            "Choosing \u03b1 = 0 . 8 , \u03b3 = 2 , \u03c7 = 2 , then solving (14) and (28) in Theorem 3 , the following solutions can be obtained:\n\u03c61 =\n( 3 . 1445\n3 . 4103\n) , \u03c62 = ( 3 . 9029\n3 . 5896\n) , \u03d5 1 = \u03d5 2 = ( 5 . 5226\n5 . 5226\n) , \u03c8\n11 11 = \u03c8 31 11 =\n( 0 . 0870\n0 . 2608\n) ,\n\u03c8 12 11 = \u03c8 32 11 =\n( 0 . 1765\n0 . 3484\n) , \u03c8\n11 12 = \u03c8 31 21 =\n( 0 . 2440\n0 . 5413\n) , \u03c8\n12 12 = \u03c8 32 21 =\n( 0 . 2267\n0 . 2460\n) , \u03c8\n21 1 =\n( 0 . 0509\n0 . 1458\n) ,\n\u03c8 11 21 = \u03c8 31 12 =\n( 0 . 0860\n0 . 2591\n) , \u03c8\n12 21 = \u03c8 32 12 =\n( 0 . 1645\n0 . 3344\n) , \u03c8\n11 22 = \u03c8 31 22 =\n( 0 . 2413\n0 . 5438\n) , \u03c8\n22 1 =\n( 0 . 0989\n0 . 2086\n) ,\n\u03c8 12 22 = \u03c8 32 22 =\n( 0 . 2229\n0 . 2447\n) , \u03c8\n21 2 =\n( 0 . 1383\n0 . 3334\n) , \u03c8\n22 2 =\n( 0 . 1273\n0 . 1392\n) ,\n\u03c8 51 1 = \u03c8 52 1 = \u03c8 51 2 = \u03c8 52 2 =\n( 0 . 9016\n0 . 9016\n) , \u03c8\n4 p i 1 i 2 = \u03c8 6 p i 1 i 2\n\u2261 ( 1 . 1722\n1 . 1722\n) for i 1 , i 2 , p \u2208 { 1 , 2 } .\nTake \u02c6 v1 1 = \u0302 v1 2 = \u0302 v2 1 = \u0302 v2 2 = (0 . 2 0 . 1) T and \u02dc v1 1 = \u0303 v1 2 = \u0303 v2 1 = \u0303 v2 2 = (0 . 1 0 . 1) T , then from the above results and (39) , one gets\n\u02c6 A 1 1 = \u02c6 A 1 2 = \u02c6 A 2 1 = \u02c6 A 2 2 =\n( 0 . 0816 0 . 0816\n0 . 0816 0 . 0816\n) , \u02c6 B\n1 1 =\n( 0 . 0046 0 . 0132\n0 . 0046 0 . 0132\n) ,\n\u02c6 B 1 2 =\n( 0 . 0125 0 . 0302\n0 . 0125 0 . 0302\n) , \u02c6 B\n2 1 =\n( 0 . 0090 0 . 0189\n0 . 0090 0 . 0189\n) , \u02c6 B\n2 2 =\n( 0 . 0115 0 . 0126\n0 . 0115 0 . 0126\n) ,\n\u02c6 C 1 1 = \u02c6 C 1 2 =\n( 1 . 7333 1 . 7333 ) , \u02c6 C\n2 1 = \u02c6 C 2 2 =\n( 0 . 7517 0 . 7517 ) , \u02c6 D\n1 1 =\n( 0 . 2258 0 . 6769 ) ,\n\u02c6 D 2 1 =\n( 0 . 1783 0 . 3216 ) , \u02c6 D\n1 2 =\n( 0 . 2232 0 . 6725 ) , \u02c6 D\n2 2 =\n( 0 . 1662 0 . 3199 ) .\nHence, the controller gain matrices can be designed as\n\u02c6 A 1 1 = \u02c6 A 1 2 = \u02c6 A 2 1 = \u02c6 A 2 2 =\n( 0 . 0816 0 . 0816\n0 . 0816 0 . 0816\n) , \u02c6 B\n1 1 =\n( 0 . 0046 0 . 0132\n0 . 0046 0 . 0132\n) ,\n\u02c6 B 1 2 =\n( 0 . 0125 0 . 0302\n0 . 0125 0 . 0302\n) , \u02c6 B\n2 1 =\n( 0 . 0090 0 . 0189\n0 . 0090 0 . 0189\n) , \u02c6 B\n2 2 =\n( 0 . 0115 0 . 0126\n0 . 0115 0 . 0126\n) ,\n\u02c6 C 1 1 = \u02c6 C 1 2 =\n( 0 . 3467 0 . 3467\n0 . 1733 0 . 1733\n) , \u02c6 C 2 1 = \u02c6 C 2 2 = ( 0 . 1503 0 . 1503\n0 . 0752 0 . 0752\n) , \u02c6 D\n1 1 =\n( 0 . 0226 0 . 0677\n0 . 0226 0 . 0677\n) ,\n\u02c6 D 2 1 =\n( 0 . 0178 0 . 0322\n0 . 0178 0 . 0322\n) , \u02c6 D\n1 2 =\n( 0 . 0223 0 . 0673\n0 . 0223 0 . 0673\n) , \u02c6 D\n2 2 =\n( 0 . 0166 0 . 0320\n0 . 0166 0 . 0320\n) .\nObviously, the ADT should satisfy \u03c4a > \u03c4 \u2217 a = \u2212 ln \u03c7/ ln \u03b1 = 3 . 1063 according to (13) . We choose \u03c4a = 4 in this example.\nFor simulation aim, the nonnegative boundary condition in this example is chosen as {\nx\u0304 h (0 , l) = ( 0 . 1 1 + sin (3 l) )T = x\u0304 v (l, 0) , 0 \u2264 l \u2264 20\nx\u0304 h (0 , l) = 0 = x\u0304 v (k, 0) , l, k > 20\n(46)\nand the disturbance input w (k, l) = 2 e \u22123(k + l) . The corresponding simulation results for the closed-loop system are pre-\nsented by Figs. 6\u20139 under the boundary condition (46) , where the switching signal is taken as shown in Fig. 10 . From Figs. 6\u20139 , one can see that the resulting closed-loop system is exponentially stable. Furthermore, one can obtain\u2211 \u221e\nk =0\n\u2211 \u221e\nl=0 0 . 8 k + l \u2016 z(k, l) \u2016 1 . = 1 . 0697 < 2\n\u2211 \u221e\nk =0\n\u2211 \u221e\nl=0 \u2016 w (k, l) \u2016 1 . = 4 . 4302 , i.e., the resulting closed-loop system has l 1 -gain\nbound 2.\n6. Conclusions\nIn this paper, a class of Roesser type nonlinear system has been investigated under the switched mechanism. By resorting\nto the T-S fuzzy rules, dynamic output-feedback controllers have been designed. Sufficient conditions have been presented\nunder which the resulting closed-loop system is exponentially stable and has the l 1 -gain bound \u03b3 . In addition, the criteria obtained in this paper are presented in the form of matrix inequalities, which are computationally tractable with MAT-\nLAB toolbox. Finally, from the numerical examples given in Section 5 , one can see that the approaches proposed here are\neffective. In the near future, we will focus on the stability analysis and synthesis for the positive 2-D singular systems.",
            "do m\ne\nAcknowledgements\nThis work was supported in part by the National Natural Science Foundation of China under Grant 61673110 , the Six\nTalent Peaks Project for the High Level Personnel from the Jiangsu Province of China under Grant 2015-DZXX-003, the Sci-\nentific Research Foundations of Graduate School of Southeast University YBJJ1719, and the Graduate Research and Innovation\nProgram of Jiangsu Province KYCX17_0039.\nReferences\n[1] M. Antonelli , D. Bernardo , H. Hagras , F. Marcelloni , Multiobjective evolutionary optimization of type-2 fuzzy rule-based systems for financial data classification, IEEE Trans. Fuzzy Syst. 25 (2) (2017) 249\u2013264 . [2] L. Benvenuti , A.D. Santis , L. Farina , Positive Systems, Springer-Verlag, Berlin, Heidelberg, 2003 . [3] A. Benzaouia , A.E. Hajjaji , Advanced Takagi-Sugeno Fuzzy Systems: Delay and Saturation, Springer International Publishing Switzerland, 2014 . [4] A . Benzaouia , A . Hmamed , F. Tadeo , A .E. Hajjaji , Stabilisation of discrete 2D time switching systems by state feedback control, Int. J. Syst. Sci. 42 (3)\n(2011) 479\u2013487 . [5] H. Chen , J. Liang , Z. Wang , Pinning controllability of autonomous boolean control networks, Sci. China Inf. Sci. 59 (7) (2016) 070107 . [6] X. Chen , M. Chen , W. Qi , J. Shen , Dynamic output-feedback control for continuous-time interval positive systems under L 1 performance, Appl. Math. Comput. 289 (2016) 48\u201359 . [7] X. Chen , J. Lam , H.-K. Lam , Positive filtering for positive Takagi-Sugeno fuzzy systems under l 1 performance, Inf. Sci. 299 (2015) 32\u201341 . [8] R. Cimochowski , Asymptotic stability of the positive switched 2D linear systems described by the Roesser models, in: Proceedings of the 16th Inter-\nnational Conference on Methods and Models in Automation and Robotics (MMAR), 2011, pp. 402\u2013406 .\n[9] G.S. Deaecto , Dynamic output feedback H \u221e control of continuous-time switched affine systems, Automatica 71 (2016) 44\u201349 . [10] C. Du , L. Xie , H \u221e Control and Filtering of Two-Dimensional Systems, Springer-Verlag, Berlin, Heidelberg, 2002 . [11] Z. Duan , Z. Xiang , H.R. Karimi , Delay-dependent exponential stabilization of positive 2D switched state-delayed systems in the Roesser model, Inf. Sci. 272 (2014) 173\u2013184 . [12] N.H. El-Farra , P. Mhaskar , P.D. Christofides , Output feedback control of switched nonlinear systems using multiple Lyapunov functions, Syst. Control Lett. 54 (12) (2005) 1163\u20131182 . [13] G. Feng , Analysis and Synthesis of Fuzzy Control Systems: A Model-Based Approach, CRC Press, Taylor & Francis Group, Boca Raton, 2010 . [14] M. Gao , L. Sheng , Y. Liu , Robust H \u221e control for T-S fuzzy systems subject to missing measurements with uncertain missing probabilities, Neurocomputing 193 (2016) 235\u2013241 . [15] F. Gaxiola , P. Melin , F. Valdez , J.R. Castro , O. Castillo , Optimization of type-2 fuzzy weights in backpropagation learning for neural networks using GAs and PSO, Appl. Soft Comput. 38 (2016) 860\u2013871 . [16] W.M. Haddad , V. Chellaboina , Q. Hui , Nonnegative and Compartmental Dynamical Systems, Princeton University Press, Princeton, NJ, 2010 . [17] T. Kaczorek , Positive 1D and 2D Systems, Springer-Verlag, London, 2002 . [18] A .A . Kaddour , E.H.E. Mazoudi , K. Benjelloun , N. Elalami , Static output-feedback controller design for a fish population system, Appl. Soft Comput. 29\n(2015) 280\u2013287 . [19] O.M. Kwon , M.J. Park , J.H. Park , S.M. Lee , Stability and stabilization of T-S fuzzy systems with time-varying delays via augmented Lyapunov\u2013Krasovskii\nfunctionals, Inf. Sci. 372 (2016) 1\u201315 . [20] L. Li , W. Wang , Fuzzy modeling and H \u221e control for general 2D nonlinear systems, Fuzzy Sets Syst. 207 (2012) 1\u201326 . [21] S. Li , Z. Xiang , H. Lin , H.R. Karimi , State estimation on positive Markovian jump systems with time-varying delay and uncertain transition probabilities, Inf. Sci. 369 (2016) 251\u2013266 . [22] Z. Li , H. Gao , H.R. Karimi , Stability analysis and H \u221e controller synthesis of discrete-time switched systems with time delay, Syst. Control Lett. 66 (2014)\n85\u201393 . [23] D. Liberzon , Switching in Systems and Control, Birkh\u00e4user Boston, 2003 . [24] X. Liu , Stability analysis of a class of nonlinear positive switched systems with delays, Nonlinear Anal. Hybrid Syst. 16 (2015) 1\u201312 . [25] P. Melin , O. Castillo , A review on type-2 fuzzy logic applications in clustering, classification and pattern recognition, Appl. Soft Comput. 21 (2014)\n568\u2013577 . [26] G.M. M\u00e9ndez , M.A. Hern\u00e1ndez , Hybrid learning mechanism for interval A2-C1 type-2 non-singleton type-2 Takagi-Sugeno-Kang fuzzy logic systems,\nInf. Sci. 220 (2013) 149\u2013169 .\n[27] B. Niu , J. Zhao , Stabilization and L 2 -gain analysis for a class of cascade switched nonlinear systems: an average dwell-time method, Nonlinear Anal. Hybrid Syst. 5 (4) (2011) 671\u2013680 . [28] M. Philippe , R. Essick , G.E. Dullerud , R.M. Jungers , Stability of discrete-time switching systems with constrained switching sequences, Automatica 72 (2016) 242\u2013250 . [29] J. Shen , J. Lam , Static output-feedback stabilization with optimal L 1 -gain for positive linear systems, Automatica 63 (2016) 248\u2013253 . [30] P. Shi , Y. Xia , G.P. Liu , D. Rees , On designing of sliding-mode control for stochastic jump systems, IEEE Trans. Autom. Control 51 (1) (2006) 97\u2013103 ."
        ]
    },
    {
        "image_filename": "designv11_13_0003705_ecce.2018.8558256-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003705_ecce.2018.8558256-Figure8-1.png",
        "caption": "Fig. 8. A sketch of d and q flux lines, related to currents imposed in the on-load simulation. Along the d\u2212axis the stator and rotor flux is due only to isd. Along the q\u2212axis stator and rotor currents are equal and opposite, so that only leakage flux are produced by isq and irq .",
        "texts": [
            " Actually, the first simulation of the procedure, presented in Section IV, is enough for the IM features prediction. The current vectors are imposed in the simulation as in (18). It is important to notice that, along the d-axis, only the stator current isd is imposed, in order to create the main flux linkages \u03bbsd and \u03bbrd. For convenience, the rotor current ird is set equal to zero, as in the RFO model. Along the q-axis stator and rotor torque currents are equal and opposite, producing only leakage fluxes, as shown in Fig. 8. From the field solution, stator and rotor dq flux linkages are computed and the machine inductances are derived using (21). The mutual inductance M and the rotor self inductance Lr = M + Llr are achieved taking into account the on-load iron saturation, when, besides the flux current isd, even q-axis torque currents are imposed. In particular, the cross saturation between d and q axes magnetic paths and the effect of on-load stator space harmonics are considered. In the postprocessing, the FEA results have to be processed using the RFO model equations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure11-1.png",
        "caption": "Fig. 11. Undercutting \u2014 type IIa: z=6; x=xmin=0.449; \u03b4r\u204e=0; \u03b4t\u204e=0.045; \u03b1=20\u00b0; ha\u204e=0.8; c*=0.7; \u03c1*=3.2.",
        "texts": [
            " 10 it is immediately seen that if the rack-cutter parameters are \u03b1=20\u00b0, ha\u204e=1, c*=0.25, \u03c1*=0.38, when cutting a gear of parameters z=6 and x=\u22120.2 an undercutting of type I is obtained, because the tip line g\u2212g of the rectilinear profile of the rack-cutter does not cross the meshing line in the part A'PK. In this case, the teeth are undercut simultaneously in a radial and tangential direction and the undercutting indices are respectively: \u03b4r=1.25 mm; \u03bbr=12.74%; \u03b4t=1.47 mm; \u03bbt=20.66%. The presence of anundercutting of type II\u0430 is found from Fig. 11,where the generation of the gear profiles (z=6; x=xmin=0.449), is realized by the rack-cutter of parameters: \u03b1=20\u00b0; ha\u204e=0.8; c*=0.7; \u03c1*=3.2. It is seen from the figure that the teeth are not undercut of type I, as the tip line g\u2212g at x=xmin passes through point A'. Besides, the starting point b of the involute profile ba lies on the base circle, and as a result, the teeth are not undercut in a radial direction (\u03b4r=0 mm, \u03bbr=0%). In this case only a tangential undercutting is obtained, caused by the rack-cutter fillet AF, where \u03b4t=0.47 mm and \u03bbt=4.82%. This is determined by the fact that the rack-cutter fillet (profiled over a circle of a radius \u03c1*>\u03c1max\u204e=2.05) is placed in the area ADE (Fig. 7), outside the boundary area ACE. When generating the teeth, shown on Fig. 12, the parameters of the gear and rack-cutter, excluding the coefficient \u03c1*=10, are the same as on Fig. 11. And in this case, from Fig. 12 it is seen that the teeth are not undercut of type I, as the tip line g\u2212g passes through point A'. The undercutting obtained is of type IIb and is also provoked by the rack-cutter fillet AF. In this case, due to the larger value of \u03c1*, the fillet curve AF is positioned outside both boundary areas ACE and ADE. As a result, the teeth are undercut in a tangential, as well as in a radial direction and the undercutting indices are respectively: \u03b4r=0.68 mm; \u03bbr=4.76%; \u03b4t=1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002066_2013.40949-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002066_2013.40949-Figure4-1.png",
        "caption": "FIG. 4 Oval footprint of a conventional tire compared with that of a radial-ply tire.",
        "texts": [
            " To these characteristics can be added another, less rolling resistance. On hard surfaces our company's radial-ply belted construction truck tire has up to 20 percent less rolling resistance than a regular truck tire. Lower rolling re sistance, perhaps to a different degree, should also be true of radial-ply tractor tires. Another result of greater deflection of the radial plies and the stiffness of 1962 \u2022 TRANSACTIONS OF THE ASAE 109 the beltlike breaker is an increase in ground contact area. Fig. 4 compares the oval footprint of a conventional tire with the longer, more rectangular footprint of a radial-ply tire. The in crease in contact area should provide more flotation and less soil compaction with radial-ply tires. A result of greater deflection under load is a lower spring rate for radial-ply tires. Lack of lateral stability has already been touched on, along with some of the items that could improve it. A radial-ply belted tire has less lateral stability than conventional tires. This could affect tractor and implement per formance in precision farming opera tions, especially along the side of a hill"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003255_3132446.3134888-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003255_3132446.3134888-Figure1-1.png",
        "caption": "Figure 1: Kinematic model of 2-dimensonal robot",
        "texts": [
            " Restart procedure with different initial guesses is incorporated in the algorithm to avoid the problem of getting IK solution at local minima. This avoids the use of evolutionary optimization techniques, which are computationally expensive. The proposed approach can be easily incorporated to real-time applications because of its efficiency. Effectiveness of the method is demonstrated by simulating various case studies of redundant manipulators, shown in Section 6. Consider a \ud835\udc5b \u2212 \ud835\udc37\ud835\udc42\ud835\udc39 planar manipulator shown in the fig 1. If the length of \ud835\udc56\ud835\udc61\u210elink is \ud835\udc59\ud835\udc56, \ud835\udf03\ud835\udc56 represents the angle between the \ud835\udc56\ud835\udc61\u210e link and the \ud835\udc65 \u2212 \ud835\udc4e\ud835\udc65\ud835\udc56\ud835\udc60. The end-effector position \ud835\udc38(\ud835\udc65\ud835\udc61\ud835\udc52, \ud835\udc66\ud835\udc61\ud835\udc52) is given by \ud835\udc65\ud835\udc61\ud835\udc52 = \ud835\udc591 cos \ud835\udf031 + \ud835\udc592 cos \ud835\udf032 + \ud835\udc593 cos \ud835\udf033 + \u2026 . +\ud835\udc59\ud835\udc5b cos \ud835\udf03\ud835\udc5b (1) \ud835\udc66\ud835\udc61\ud835\udc52 = \ud835\udc591 sin \ud835\udf031 + \ud835\udc592 sin \ud835\udf032 + \ud835\udc593 sin \ud835\udf033 + \u2026 . +\ud835\udc59\ud835\udc5b sin \ud835\udf03\ud835\udc5b (2) Forward kinematic model of the end effector is related as \ud835\udc38(\ud835\udc65\ud835\udc61\ud835\udc52, \ud835\udc66\ud835\udc61\ud835\udc52)=\ud835\udc53(\ud835\udf03), \ud835\udc38 \u2208 \ud835\udc45\ud835\udc5a, \ud835\udf03\ud835\udf16\ud835\udc45\ud835\udc5b (3) where \ud835\udc5a is the task space dimension and \ud835\udc5b is the joint space dimension. Redundant manipulators have joint space dimension (n) greater than task space dimension (m)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003580_s11661-018-4885-8-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003580_s11661-018-4885-8-Figure2-1.png",
        "caption": "Fig. 2\u2014(a) Coupling bionic sample; (b) stress distribution of bionic surface; (c) cross-sectional sketch of laser re-melting unit.",
        "texts": [
            " Different graphite types of gray cast iron were obtained by changing the material composition and solidification method during casting. A 5-kg medium frequency induction furnace was used for melting. No. A and C were obtained by the wet sand casting with the advantages of slow cooling and long gestation time. No. B and E specimens were obtained by the resin self-hardening sand casting with a thin mold wall. No. D specimen was obtained by the metal casting. Experimental gray cast iron specimens with a size of 36 9 20 9 6 mm3 were cut using an electric spark machine. As shown in Figure 2(a), for each sample, a 2-mm diameter round hole and a 2-mm-deep notch were cut at the two ends of the specimen. The hole as shown was used to fix sample onto the plate of the thermal fatigue experimental machine (shown in Figure 3). The METALLURGICAL AND MATERIALS TRANSACTIONS A notch is used to induce stress concentration. To avoid possible experimental errors caused by the surface machining marks, the specimens were mechanically polished by grinding with sand papers and polishing with cotton cloth. The coupled surface of matrix material (soft) and the unit (hard) is named as bionic surface (shown in Figure 2). In the bionic surface, the bionic units designed with 45 deg mesh and 0 deg stripe (angles are determined by comparing orientations with longer edges) were fabricated on one side of the specimens, while the corresponding side is the nonbionic surface. The macro-distance between every two parallel ridges (mesh units) as shown is fixed at 5 mm. The unit was processed by a solid-state Nd:YAG laser of 1.06 um wavelength and a maximum power 500 W. The laser processing parameters are listed in Table II"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure22-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure22-1.png",
        "caption": "Fig. 22. Contact stress distribution of the tooth pairs without consideration of the influence of the carrier with the planet shafts.",
        "texts": [
            " Either the contact stress of the sun-planet or the annulus-planet tooth pair in Case A is distributed almost evenly along the face-width, only the concentrated stress occurs on both the face-ends due to edge effect, as the results in Figs. 20 and 21 show. As Fig. 20 clearly shows, the twist deformation of the sun gear ( Case B ) cause unevenly distributed stress of the sun-planet tooth pair, where the enlarged stress occurs near the input side and reduced stress near the output side. The trapezoid-shaped contact pattern, as illustrated in Fig. 22 (a), shows that not only the stress but also the contact area is enlarger near the input side of the sun gear. This strongly non-uniform variation of the contact stress is also caused by the reduced twist stiffness due to the hollow design of the sun gear. However, this twist deformation has no effect on the stress distribution of the annulus-planet tooth pair, that is why the calculated results in Case A and B are almost the same, Fig. 21 . The corresponding contact pattern, as shown in Fig. 22 (b), is saddle-shaped. The influence of the combined deformation of the carrier with the planet shafts can be observed with comparison of the analysis results of the annulus-planet tooth pair in Case B ( A ) and C , see Fig. 21 . The difference shows that the stiffness of the carrier on the output side is more rigid than elsewhere. This effect is opposite to that of the twist deformation of the sun gear. As a consequence, the unevenly distributed stress of the sun-planet tooth pair due to the twist of the sun gear is thus compensated by the twist of the carrier with the planet shafts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000897_jf502160d-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000897_jf502160d-Figure1-1.png",
        "caption": "Figure 1. Effects of dissolving solvents and ratios of solvent volume to the reaction mixture on the xanthine/xanthine oxidase reaction. The reaction mixture contained 6.67 \u03bcL of different solvents, 193.33 \u03bcL of substrate solution (6.43 \u03bcL of 4 mM xanthine, 12.86 \u03bcL of 4 mM NBT, and 174.04 \u03bcL of 0.05 M PBS), and 20 \u03bcL of 40 mU/mL xanthine oxidase. The absorbance at 560 nm was determined for each reaction mixture after 20 min of reaction terminated by the addition of 20 \u03bcL of 0.6 M HCl. Four different volume ratios (1:10, 1:20, 1:30, and 1:40, v/v) were tested. ACE, ACN, ET, IPA, and THF represented acetone, acetonitrile, ethanol, isopropanol, and tetrahydrofuran, respectively. Data are means of triplicate measurements. Vertical bars represent the standard deviations.",
        "texts": [
            "39 This suggests the feasible use of organic cosolvents for the hydrophilic and lipophilic antioxidant samples tested in the enzymatic assay. Therefore, several organic solvents including DMSO, acetone, acetonitrile, isopropanol, and tetrahydrofuran, as well as water, were selected to investigate the solvent effect on the reaction system. With a fixed ratio of solvent volume to the reaction mixture (1:30), none of these organic solvents had shown a significant effect on absorbance values compared with water (Figure 1), indicating the potential use of these solvents as sample dissolving solvents. This is especially important and useful for samples with lower solubility in water. In addition, three other different ratios of solvent volume to the reaction mixture (1:10, 1:20, and 1:40, v/v) were investigated and compared with the ratio of 1:30. As shown in Figure 1, no significant absorbance changes were observed among the 1:20, 1:30, and 1:40 ratios of solvent volume to reaction mixture, regardless of the nature of dissolving solvents. However, in a 1:10 ratio, all of the solvents except water and ethanol resulted in obvious decreases of the absorbance values (Figure 1). Therefore, the organic solvent effect should not be ignored in a relatively high ratio (1:10) of solvent volume to the reaction mixture (except water and ethanol). In consideration of its extensive use as a dissolving solvent in screening tests, DMSO with a 1:30 ratio of volume to the reaction mixture was used in this study. Effect of Different Terminators on the X/XO Reaction. Significant pH changes can affect enzyme activity, which might be an effective way to terminate an enzymatic reaction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003714_s12008-018-0520-6-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003714_s12008-018-0520-6-Figure7-1.png",
        "caption": "Fig. 7 Simulation with radiation heat losses only",
        "texts": [],
        "surrounding_texts": [
            "The developed model was first utilized to examine the effect of convection heat transfer during theAMprocess. The simulation analysiswas carried out for two cases,with andwithout convection heat losses, while all other process parameters were kept unchanged. Figures 6 and 7 illustrate the effect of convective boundary condition on the temperature variation within the part for a laser speed of \u03c0 mm/s. The predictions are obtained in each case for the time steps after activating the layers 1, 2, 3 and 4 at the time instant, when the laser beam reaches the middle of the corresponding layer. The results indicated that the absolute values of the temperature difference between the two cases increase with time. The computed temperature differences lie in the interval between 15 and 30 \u00b0C and reach a maximum relative value of approximately 3%. Subsequently, neglecting the convection losses would not lead to significant errors. Figure 8 shows the temperature variation along the height of the part for both cases. This figure clearly shows the insignificance of convective boundary condition on the results (h 30 W/m2 K). Figure 9 illustrates the thermal cycles at points A, B and C corresponding to the substrate and the two deposited layers 2 and 4, respectively. The peaks indicate the time instants at which the laser beam passes through the pre-defined locations.At themid-point of the 2nd layer, the initial temperature peak is approximately 1750 \u00b0C. After 1.5 s, the heat is conducted away and the temperature decreases to a value of about 540 \u00b0C. The mid-point of layer 4 has experienced a similar thermal cycle, where the maximum temperature has a value of 2113 \u00b0C. The temperature peaks at points B and C indicate that the melting point of the material has been reached. An interesting observation is noticed in Fig. 8 that after the deposition of the fourth layer, the mid-point temperature of the 2nd layer reached 1650 \u00b0C. This aging and tempering effects may affect the mechanical properties of the part including the residual stresses and its strength. Figure 10 demonstrates the effect of varying the speed of the laser beam on the temperature distribution of the part after completing building the last layer (4th layer). It can be seen from Fig. 10 that as the velocity of the laser beam increases, the temperature of the part decreases comparedwith the case of lower velocity. This is can be attributed to less time of exposure to the laser beam. Figure 11 shows the temperature variation along the height of the part for various laser travel velocities. The plotted data was taken at the time instant after the 4th layer was deposited. The corresponding maximum temperature were about 2105 \u00b0C, 2010 \u00b0C, 1700 \u00b0C and 1510 \u00b0C for tested scan speed of \u03c0 mm/s, 2\u03c0 mm/s, 4\u03c0 mm/s and 6\u03c0 mm/s, respectively. Figure 10 clearly shows that the temperature along the height of the part is the lowest for the higher velocity (6\u03c0 mm/s) and this is due to the shorter interaction time between the laser beam and the material layer during the scanning process. However, the lowest travel velocity exhibits themaximum temperature."
        ]
    },
    {
        "image_filename": "designv11_13_0003068_iceee2.2017.7935821-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003068_iceee2.2017.7935821-Figure1-1.png",
        "caption": "Figure 1. Quadcopter forces and attitude angles",
        "texts": [
            "81 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 B= 0 0 0 0 (4) 0 0 0 0 0 0 0 0 0 . 5917 0 0 0 0 36.7769 0 0 0 0 36. 7769 0 0 0 0 18.9415 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 C=O 0 0 0 0 1 0 0 0 0 0 0 (5) 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 In Figure 1, angles and forces of the quadcopter is presented, also in Figure 2, hardware system configuration of the modelled quadcopter is seen. Totally, there are 12 states of the modelled quadcopter system which are (x, y, Z, qJ, 8, 1./1, X, y, i, <jJ, (), l/J). x, y, z are positions along each associated axis. (qJ, 8, 1./1) are angles relative to body frame and following six states are derivatives, respectively. Note that, since there is not any feedforward term in dynamics, D is a zero matrix. There are 4 controllable states in this model while all of the states are observable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000940_ems.2015.85-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000940_ems.2015.85-Figure1-1.png",
        "caption": "Figure 1. Earth-fixed (xI , yI , zI) and body-fixed (xB , yB , zB) frames [1].",
        "texts": [
            "85 313 The surface vessel model has 6-DOF: surge, sway, yaw, heave, roll, and pitch, which can be simplified to motion in surge, sway, and yaw under the following assumptions [1]: 1) The surge, sway, and yaw induced by wind and currents are negligible, 2) The inertia, added mass, and hydrodynamic damping matrices are diagonal, 3) The available control variables are surge force and yaw moment. Based on that, the 3-DOF model will be [13]: x\u0307 = u cos(\u03c8)\u2212 \u03c5 sin(\u03c8) y\u0307 = u sin(\u03c8) + \u03c5 cos(\u03c8) \u03c8\u0307 = r u\u0307 = m2 m1 \u03c5r \u2212 d1 m1 u+ 1 m1 \u03c4u \u03c5\u0307 = m1 m2 ur \u2212 d2 m2 \u03c5 r\u0307 = (m1 \u2212m2) m3 u\u03c5 \u2212 d3 m3 u+ 1 m3 \u03c4r. (1) Here, x and y are the positions, and \u03c8 is the heading angle of the ship with respect to the earth-fixed frame, u and \u03c5 are linear velocities and r is the angular velocity with respect to body-fixed frame (see Figure1), the parameters m1,m2,m3 are the ship inertia including added mass effects, and d1, d2, d3 are the hydrodynamic damping coefficients. Input constraint is imposed to account for the maximum surge force and yaw moment of the ship. A time-varying reference trajectory is generated by a virtual ship with the same dynamics as (1): x\u0307r = ur cos(\u03c8r)\u2212 \u03c5r sin(\u03c8r) y\u0307r = ur sin(\u03c8r) + \u03c5 cos(\u03c8r) \u03c8\u0307r = rr u\u0307r = m2 m1 \u03c5rrr \u2212 d1 m1 ur + 1 m1 \u03c4ur \u03c5\u0307r = m1 m2 urrr \u2212 d2 m2 \u03c5r r\u0307r = (m1 \u2212m2) m3 ur\u03c5r \u2212 d3 m3 ur + 1 m3 \u03c4rr "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000747_1077546313483786-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000747_1077546313483786-Figure1-1.png",
        "caption": "Figure 1. Diagram for steering system with clearance in steering linkage mechanism.",
        "texts": [
            " Because it is still difficult for numerical analysis of a high-dimension nonlinear system with multi-clearances, only one clearance is considered in the dynamic modeling of the system, and without loss of generality, the system with clearance in the kinematic pair between the crank and link is employed as an example to discuss the dynamic response of the system. The diagram for the steering system with clearance in the kinematic pair between the steering tie rod and left trapezoidal arm of a certain vehicle with dependent suspension is shown in Figure 1, where (a) is the view from top, and (b) is the view from behind. The dynamic model of the vehicle shimmy system with clearance in steering linkage mechanism can be obtained based on the following assumptions. 1. The vehicle is driven in line with the speed of v, and the roughness of the road is neglected. 2. Only the dynamic response of the vehicle due to lateral excitation is considered. 3. The steering wheel is fixed in the position of directing ahead in line. 4. Other clearances in the steering system are neglected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003333_s10845-018-1391-1-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003333_s10845-018-1391-1-Figure7-1.png",
        "caption": "Fig. 7 Heat source model",
        "texts": [
            " The boundary condition can be expressed by kn \u2202T \u2202n \u2212 q + h (T \u2212 T0) + \u03c3\u03b5 ( T 4 \u2212 T 4 0 ) = 0 (x, y, z) \u2208 S t > 0 (24) where S is the boundary to be calculated, kn is the thermal conductivity on the boundary surface S, h is the convection coefficient, \u03b5 is the radiation heat transfer coefficient, and q is the heat flux on the boundary surface S. Typically, the complex physics of heat generation or weld pool is simplified considerably and replaced by a heat input model. In this work, the previously developed body heat source model, made by combining double-ellipsoids, rotating-Gaussian, and a cone (Jiang et al. 2016), is adopted for FEMsimulation. The illustration of the heat sourcemodel is plotted in Fig. 7. Thermo-physical properties of the material will change as the temperature increases during the welding process, especially when the temperature between liquidus and solidus (Islam et al. 2014). Therefore, thermo-physical properties of stainless steel need to be set in the finite element simulation. The specific heat capacity and thermal conductivity change with temperature are listed in Table 2. After calculating the temperature field of fiber laser keyhole welding, the thermal analysis units of the model are transformed into the corresponding structural units"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001536_s00170-018-2494-8-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001536_s00170-018-2494-8-Figure1-1.png",
        "caption": "Fig. 1 General arrangement of the welding torch, wire feed and inert trailing shielding used to produce additively manufactured samples",
        "texts": [
            " The combined torch and wire feed arrangement was mounted above a linear actuator, with height adjustment possible through a manually operated rack-and-pinion. Inert gas shielding of the weld zone was achieved with welding grade pure argon using pre- and post-flow options. Additionally, a custom fabricated trailing shield was used to ensure that a sufficiently large shielding envelope was generated to the rear of the fusion zone to prevent post-weld atmospheric contamination of the successive weld beads. The general arrangement of GTAW equipment is illustrated in Fig. 1. Shielding gas flow rates and other common process parameters are detailed in Table 2. Side milling was carried out on the GTAW additively manufactured Ti-6Al-4V multi-layered deposited structures. A total of four build-up structures were produced using the process variables listed in Table 3. All structures consisted of a single weld pass per layer, producing a wall width of approximately 8 mm (schematic shown in Fig. 2a). For each sample an initial \u2018skimming\u2019milling pass was performed to generate a smooth consistent surface such that the thickness of the wall was machined down to ~ 7 mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000108_j.procir.2012.10.015-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000108_j.procir.2012.10.015-Figure1-1.png",
        "caption": "Fig. 1. Concept of milling-combined laser sintering system",
        "texts": [
            " Eiji Shamoto hardness, thermal conductivity, residual stress and deformation were also measured. After that, high speed milling was performed to investigate the machinability of the sintered materials. Cutting force and cutting temperature of both sintered materials were experimentally measured using ball end mill with 6 mm of diameter. The cutting temperature was measured by using a three-color pyrometer which was developed by one of the authors[3]. Bulk carbon steel (JIS S55C) was selected as reference steel. Concept of milling-combined laser sintering system is illustrated in Fig.1. The system consists of two alternating processes, which are forming sintered part by selective laser melting (SLM) and high speed milling for surface finishing. By using CAD, a 3D model was divided into sliced layers and transferred to the MLSS. Before the sintering process started, a sandblasted steel base plate was placed on the forming table. The powder table was lifted up while the forming table was moved down respectively, and the predetermined layer thickness of metal powder was deposited on a base plate by the recoater blade",
            " After forming a few layers of sintered material, the high speed milling process is executed at the periphery surface. These processes were repeated until a complete model was created and were performed in a nitrogen atmosphere at room temperature to prevent oxidization. In this paper, two different types of sintered materials; Chromium molybdenum steel (SCM) sintered material and maraging steel (MAS) sintered material were made by using selective laser melting (SLM) method. The sintered materials were fabricated using different types of metal powder. Fig. 1 shows scanning electron images of (a) SCM powder and (b) MAS powder, respectively. Characteristics of both metal powders are shown in Table 1. The mean particle size of both metal powders is 30 \u03bcm. Thermal conductivity of SCM powder is 9 W/m\u00b7K and MAS powder is 17 W/m\u00b7K, respectively, which were measured using a technique developed by the author[4]. Sintered materials were made by the MLSS without performing the milling process for purpose of cutting force and cutting temperature measurement. The characteristics of the sintered materials were summarized in Table 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001523_s40194-018-0625-3-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001523_s40194-018-0625-3-Figure4-1.png",
        "caption": "Fig. 4 Experimental setup\u2014oscillated fillet welding of lap joint",
        "texts": [
            " The spatter behaviour was analysed by high speed videos. The surface appearance was assessed by macroscopic and microscopic pictures. The main criterion hereby was a continuously closed fillet weld without pores or lack of fusion. In addition no full penetration welds were allowed. Here zinc coated material of type H260LAD + Z100 was used, a degassing gap of 0.2 mm height was set by a gauge plate to avoid unwanted back coupling effects caused by zinc vapour. The basematerials\u2019 chemical composition is listed in Table 1. Figure 4 shows the setup of the fillet weld in the lap joint experiments. Due to industrial demands, the trials were conducted at a welding speed of 2.5 m/min, a constant lateral angle (\u03b1) of 10\u00b0 and drag angle (\u03b2) of 0\u00b0. The oscillation was applied transversal (y-direction) to the welding direction with a sinusoidal oscillation profile. The focal position was set to the upper sheets surface; the laser power was adjusted between 1500\u20131700 W. The bead-on-plate welding trials showed a clear increase of spatter formation for oscillated welds compared to a conventional straight-line welding process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003703_j.jelechem.2018.12.001-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003703_j.jelechem.2018.12.001-Figure3-1.png",
        "caption": "Fig. 3",
        "texts": [
            " The Ag/GCN/CS ternary composites exhibited a similar peak shape and position, but the emission intensity of the Ag/GCN/CS ternary composites was far weaker than that of GCN, indicating that the doping of Ag nanoparticles and CS on the surface of GCN suppressed the recombination process of photogenerated electron-hole pairs effectively, resulting in the enhanced PEC performance of the Ag/GCN/CS ternary composites. [19, 37] The charge transfer capability was investigated by EIS. Theoretically, a smaller arc radius means that a material has a higher charge transfer rate. [11] As shown in Fig. 3a, the radius of Ag/GCN/CS/ITO was much smaller than that of Ag/GCN/ITO and the radius of Ag/GCN/ITO was smaller than that of GCN/ITO. This phenomenon demonstrates that Ag/GCN/CS/ITO possessed a higher charge transfer rate and better electrical conductivity than that of Ag/GCN/ITO and GCN/ITO. The photocurrent responses of GCN/ITO and Ag/GCN/CS/ITO are shown in Fig. 3b. Each modified unit, once the light was turned on, had an evident increased photocurrent. When the excited light disappeared, the photocurrent response instantly returned to the background value. The photocurrent response measured on Ag/GCN/CS/ITO was almost 2.5 times higher than that of GCN/ITO and almost 1.9 times higher than that of Ag/GCN/ITO, demonstrating that the synergistic effect between Ag nanoparticles and CS promoted the separation efficiency of photogenerated electron-hole pairs and transfer rate of electrons, resulting in the enhanced PEC performance of Ag/GCN/CS ternary composites",
            " Zhou, Construction of plasmonic Ag and nitrogen-doped graphene quantum dots codecorated ultrathin graphitic carbon nitride nanosheet composites with enhanced photocatalytic activity: Full spectrum response ability and mechanism insight, ACS Appl. Mater. Interfaces 9 (2017) 42816-42828. AC C EP TE D M AN U SC R IP T Fig. 1. (a) SEM image and (b) TEM image of Ag/GCN/CS ternary composites; (c) XRD spectra and (d) FT-IR of Ag/GCN/CS ternary composites and GCN, respectively (The inset: magnified XRD pattern of Ag/GCN/CS ternary composites). Fig. 2. (a) UV-vis diffuse reflectance spectra and (b) PL of Ag/GCN/CS ternary composites and GCN, respectively. Fig. 3. (a) Nyquist plots of Ag/GCN/CS/ITO, Ag/GCN/ITO and GCN/ITO in 0.1 mol L \u20131 KCl solution containing 3 mmol L \u20131 [Ru(NH3)6 3+ ]Cl3/[Ru(NH3)6 2+ ]Cl2 (1:1) mixture in the frequency range from 0.01 Hz to 10 kHz. and (b) Photocurrent responses of GCN/ITO, Ag/GCN/ITO and Ag/GCN/CS/ITO in phosphate buffer solution (0.1 mol L \u20131 , pH 7.0). Fig. 4. (a) Photocurrent responses of the Ag/GCN/CS/ITO, Ag/GCN/ITO and GCN/ITO in the absence (b, d, f) and presence (a, c, e) of 200 ng mL \u22121 4-CP; (b) Stable photocurrent responses curve of Ag/GCN/CS/ITO in the presence of 200 ng mL \u22121 4-CP; (c) Photocurrent responses at Ag/GCN/CS/ITO towards 4-CP at increasing concentrations; (d) The linear calibration curve for 4-CP detection"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002282_icarcv.2014.7064504-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002282_icarcv.2014.7064504-Figure1-1.png",
        "caption": "Figure 1. Coordinate system for berthing.",
        "texts": [
            " 1321 **)(* dCCC dorder \u2212\u2212\u2212= \u03c8\u03c8\u03c8\u03b4 \u2212=< == => 00 00 00 10,0 0,0 10,0 orderorder orderorder orderorder if \u03b4\u03b4 \u03b4\u03b4 \u03b4\u03b4 (2) where, d is desired heading angle, d1 is deviation from imaginary line, C1, C2 and C3 are coefficients. The first term of equation (2) provides the necessary correction for maintaining particular ship heading, second term belongs to minimizing yaw rate and the third term is for compensating ship\u2019s deviation from the pre-set imaginary line. After the ship successfully stops within the surrounded area of berthing goal point as shown in figure 1, the final step in to align it with actual pier. Usually, a big ship with single rudder single propeller often needs tug assistance for executing such crabbing motion. The number of tugs involves in such operation depends on size of ship as well as existing environmental disturbances. In this research, to develop a controller for side thrusts first ANN has been tried as by Tran and Im [13]. But considering wind which is mostly unpredictable, there is no other easy way to maintain consistency in teaching data which is very important to ensure effective ANN controller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001971_03008207.2015.1066779-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001971_03008207.2015.1066779-Figure6-1.png",
        "caption": "Figure 6. The normalized hysteresis area for various strains and strain-rates used in the tensile tests of articular cartilage (n\u00bc 37).",
        "texts": [
            " This is the transient modulus defined as d =d\", also referred to as apparent modulus, which differentiates from the modulus measured at equilibrium. The transient modulus is a function of loading history and so is the transient stress and strain. The hysteresis loops were plotted for the test case of 15% strain (Figure 5). The area of the loop represents the energy loss during the unloading process. For the purpose of comparison, the energy loss was normalized to the total energy at the end of the loading phase in each test to show the relative energy loss (Figure 6). The relative energy loss was more dependent on the strain-rate than strain magnitude. The loss reduced to the lowest at moderate strain-rates (around 10%/s). In addition, the energy dissipation at 40%/s and higher strain-rates was greater than that at 0.1%/s for all three strains (Figure 6). Our tensile experiments on articular cartilage showed substantial strain-rate dependence of the load response of D ow nl oa de d by [ U ni ve rs ity o f M on ta na ] at 0 1: 48 2 4 Fe br ua ry 2 01 6 the tissue on a full range of strain-rates at physiologically reasonable deformation that has not been fully examined previously. Furthermore, the strain-rate dependence was nonlinear and augmented at greater strain-magnitudes, among the three strains considered (Figure 3). The strain-rate dependence was less considerable at 3% strain",
            " This variation of non-linearity with strain-rate was different from that of cartilage in compressive testing. For example, cartilage in low strain-rate tensile testing also exhibited non-linear stress-strain relationships, in contrast with linear stress-strain relationships at nearly static compressive testing shown in experiments (19) and explained in modeling (4). The relative energy loss in tensile behavior during unloading demonstrated by the hysteresis was strongly strain-rate dependent (Figures 5 and 6). It only slightly depended on the strain magnitude (Figure 6). This trend of energy loss may be explained by a damping mechanism in the tissue. This damping mechanism caused the energy loss to be non-linear with the lowest relative energy loss at a moderate strain-rate. As the tensile loading in the tensile tests was predominantly supported by the fibers, the damping mechanism was likely provided by the fiber network, including its interaction with the fluid and proteoglycans. The hysteresis testing on single collagen fibers also showed a similar change of hysteresis with strain-rate (31)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001786_9783527344758-Figure1.10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001786_9783527344758-Figure1.10-1.png",
        "caption": "Figure 1.10 Finite element modelling of fringe field effect in a micro device.",
        "texts": [
            " (2016). As communicated in Section 1.4.1, the GMR sensors can be used, while modeling biological devices, to detect magnetic labels. When it comes to computational approach, an inverse method can be used to utilize the effect of fringe fields present on the periphery of the GMR elements, thus changing the GMR response per MNP. The study of a solution containing MNPs, flowing through microfluidic channels parallel to the GMR sensor\u2019s edge, under Poiseuille flow, can demonstrate such approach. In Figure 1.10, a schematic of the fringe field effect between two sensors is presented. For further information on the device design concept, modeling, and computations, some useful sources may be accessed (Pankhurst et al. 2003; Hamdi and Ferreira 2008; Rani 2014). With the passage of time, and with the advancement in the field of nanotechnology, more advanced and accurate algorithms (e.g. Figure 1.11) and computational models have commercially launched. The nanofluid infusion and the subsequent thermal activation of the infused NPs are two critical stages during the hyperthermia treatment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001916_icra.2015.7139346-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001916_icra.2015.7139346-Figure2-1.png",
        "caption": "Fig. 2. The experimental setup for performing needle insertions. A DC motor (Motor 1) provides the linear motion to insert the needle into the tissue. A second motor (Motor 2) attached to the needle base is able to rotate it around its axis during insertion (not used in this paper). The forces at the needle base are measured by a force/torque sensor. Images of the needle inside tissue are recorded by a Logitech C270 camera.",
        "texts": [
            " According to beam theory [10], the uniformly distributed load q1,u can again be replaced by the point load F1,u at the centroid of q1,u, which is at point B. For a uniformly distributed load, the point B is placed at 1/2 of lin. The deflection caused by F1,u is \u03b41,u = (3L\u2212 l1)l2 1 6EI F1,u (13) where l1 is defined by (3) with a = 1/2. The total tip deflection \u03b4tip,3 is given by (11) and (13): \u03b4tip,3 = \u03b41,u +\u03b42 (14) The experimental setup used to perform insertion experiments is the 2 Degree-of-Freedom (DOF) prismatic-revolute robotic system shown in Fig. 2. The needle, which represents the end-effector of the robot, can be translated along and rotated about its longitudinal axis. The translational motion is guided by a linear stage. The linear stage\u2019s carriage is coupled to a timing belt, which is driven by a DC motor (RE40, Maxon Motor AG, Sachseln, Switzerland). The rotational motion is not utilized during the experiments presented in this work. The rotational motor\u2019s shaft carries a 6 DOF force/torque transducer (50M31A3-I25, JR3 Inc., Woodland, CA, USA) to record the 2 DOF of forces and torques, which are the model inputs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure9-1.png",
        "caption": "Fig. 9. Influence of perpendicularity error on the contact region.",
        "texts": [
            " 8, it can be found that: (a) when the axial error is positive, total contact regions will deflect to the root of pinion and the topland of gear when the tooth pair conjugates from the toe to the heel, respectively; (b) the influence of axial error in TL direction gradually reduces to zero, but the influence of axial error gradually increases in TH direction; and (c) comparing with the influences of center distance error and tangential error on the contact region, the influence of axial error is close to that of tangential error, but is less than that of center distance error. (4) Influence of perpendicularity error Fig. 9 shows the deflection tendency of contact region in TL and TH directions of pinion and gear when the perpendicularity error \u03b5x = 0.03. According to Fig. 9, it can be found that: (a) when the perpendicularity error of axis holes for mounting pinion and gear is positive, total contact regions will deflect to the topland of pinion and the root of gear in the conjugation process from the toe to the heel, respectively;(b) the influence of perpendicularity error in TL direction gradually reduces to zero, but the influence of axial error in TH direction gradually increases; and (c) the influence of perpendicularity error is less than that of linear errors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000683_icra.2013.6631258-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000683_icra.2013.6631258-Figure3-1.png",
        "caption": "Fig. 3. (a) Dynamic manipulability ellipsoids (DMEs) represent the possible accelerations for each link with no prior task and (b) dynamic reconfiguration manipulability ellipsoids (DRMEs) represent the possible accelerations for intermediate links during the system executing primary task.",
        "texts": [
            " r\u0308i \u2212 J\u0307 i(q)q\u0307 = J iM \u22121[\u03c4 \u2212 h(q, q\u0307) \u2212 g(q) \u2212 Dq\u0307] (4) Here, two variables are defined as follows: \u03c4\u0303 = \u03c4 \u2212 h(q, q\u0307) \u2212 g(q) \u2212 Dq\u0307 (5) \u00a8\u0303ri = r\u0308i \u2212 J\u0307 i(q)q\u0307 = J i(q)q\u0308 (6) Then, Eq. (4) can be rewritten as \u00a8\u0303ri = J iM \u22121\u03c4\u0303 (i = 1, 2, \u00b7 \u00b7 \u00b7 , n) (7) Considering desired accelerations \u00a8\u0303rid of all links yielded by a set of joint torques \u03c4\u0303 that satisfies an Euclidean norm condition, that is, \u2016\u03c4\u0303\u2016 = (\u03c4\u03032 1 + \u03c4\u03032 2 + \u00b7 \u00b7 \u00b7 + \u03c4\u03032 n)1/2 \u2264 1, then the each tip acceleration shapes an ellipsoid in range space of J i. These ellipsoids of each link have been known as \u201cDynamic Manipulability Ellipsoid (DME)\u201d [12] (Fig.3(a)) which are described as \u00a8\u0303rid T[J i(MTM)\u22121JT i ]+ \u00a8\u0303rid \u2264 1, and \u00a8\u0303rid \u2208 R(J i) (8) where, R(J i) represents range space of J i. B. Dynamic Reconfiguration Manipulability Here we assume that the desired end-effector\u2019s acceleration r\u0308nd are given as primary task. Relation between \u00a8\u0303rn and \u03c4\u0303 is denoted to give i = n into Eq. (7), then, \u00a8\u0303rn = JnM\u22121\u03c4\u0303 (9) Solving Eq. (9) for \u03c4\u0303 yielding desired acceleration \u00a8\u0303rnd \u03c4\u0303 = (JnM\u22121)+ \u00a8\u0303rnd +[In \u2212 (JnM\u22121)+(JnM\u22121)] 1l (10) 1l is an arbitrary vector satisfying 1l \u2208 Rn",
            " (18) represents an ellipsoid expanding in m-dimensional space, holding \u22061r\u0308jd = 1\u039bj 1\u039b+ j \u22061r\u0308jd, \u22061r\u0308jd \u2208 Rm (19) which indicates that \u22061r\u0308jd can be arbitrarily generated in m-dimensional space and Eq. (16) always has the solution 1l corresponding to all \u22061r\u0308jd \u2208 Rm. On the other hand, if rank(1\u039bj) = r < m, \u2206r\u0308jd does not value arbitrarily in Rm. In this case, reduced \u2206r\u0308jd is denoted as \u22061r\u0308\u2217 jd. Then Eq. (18) is written as (\u22061r\u0308\u2217 jd)T(1\u039b+ j )T 1\u039b+ j \u22061r\u0308\u2217 jd \u2264 1 (\u22061r\u0308\u2217 jd = 1\u039bj 1\u039b+ j \u22061r\u0308jd) (20) Equation (20) describes an ellipsoid expanded in rdimensional space. These ellipsoids of Eqs. (18) and (20) are shown in Fig.3(b). C. Dynamic Reconfiguration Manipulability Shape Index(DRMSI) In this section, we propose the index evaluating DRM. Thus, by applying the singular value decomposition for this matrix \u039b , we get 1\u039bj =1 U1 j\u03a3 1 jV T j (21) 1\u03a3j = \u23a1 \u23a2\u23a2\u23a2\u23a3 r n \u2212 r 1\u03c3j,1 0 r . . . 0 0 1\u03c3j,r m \u2212 r 0 0 \u23a4 \u23a5\u23a5\u23a5\u23a6 (22) In Eqs. (21) and (22), 1U \u2208 Rm\u00d7m,1 V \u2208 Rn\u00d7n are orthogonal matrixes, and r denotes the number of non-zero singular values of 1\u039bj and \u03c3j,1\u2265 \u00b7 \u00b7 \u00b7 \u2265\u03c3j,r > 0. In addition, r\u2264m because rank(1\u039bj)\u2264m. So, dynamic reconfiguration capability of j-th link when hand of manipulator operating task can be described by following equation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001072_s00170-017-0463-2-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001072_s00170-017-0463-2-Figure3-1.png",
        "caption": "Fig. 3 Relationship between the honing wheel and the workpiece",
        "texts": [
            " The model of the unmodified workpiece is rg(ug, \u03b8g), in which ug and \u03b8g are the parameters of the flank. If the modification value is \u03be(ug, \u03b8g), the model of the modified flank could be expressed as rmg ug; \u03b8g \u00bc rg ug; \u03b8g \u00fe \u03be ug; \u03b8g \u00f01\u00de The normal vector is nmg ug; \u03b8g \u00bc \u2202rmg \u2202ug \u2202rmg \u2202\u03b8g \u00f02\u00de The shape of the honing wheel is a ZI-worm without any modification. Its model is rc(uc, \u03b8c), in which uc and \u03b8c are the parameters of the flank. The normal vector of the worm is nc uc; \u03b8c\u00f0 \u00de \u00bc \u2202rc \u2202uc \u2202rc \u2202uc. In Fig. 3, if we want to use Tooth Contact Analysis (TCA) method [19, 20], rc(uc, \u03b8c) and nc(uc, \u03b8c) should be changed to the coordinate system Sg (xg, yg, zg). \u03c6c and \u03c6g are the rotation angle of the honing wheel and the gear. rcg uc; \u03b8c\u00f0 \u00de \u00bc Mgcrc uc; \u03b8c\u00f0 \u00de ncg uc; \u03b8c\u00f0 \u00de \u00bc Mgcnc uc; \u03b8c\u00f0 \u00de Mgc is the transfer matrix from coordinate system Sc (xc, yc, zc) to Sg (xg, yg, zg). At contact point M, Eq. (3) could be acquired according to TCA method. rcg uc; \u03b8c;\u00f0 \u00de \u00bc rmg ug; \u03b8g nmg ug; \u03b8g ncg uc; \u03b8c\u00f0 \u00de \u00bc 0 \u00f03\u00de Equation (3) could deduce the coordinates of the contact point of the wheel and the workpiece, and the rotation angles of them when they rotate to the point M",
            " Actually, the honing wheel moves along the axis of the gear. When the honing wheel process the gear at each axial position of the gear, there will be a function of the transmission ratio i(\u03c6c) at each position. If the axial position is defined as l, the function of the transmission ratio of the whole process will be expressed as i = i(\u03c6c, l). During the gear honing process, the theoretical rotation speed \u03c9c is constant. Then the rotation speed of the workpiece will be computed as \u03c9g = i(\u03c6c, l)\u03c9c based on the transmission ratio acquired previously. In Fig. 3, there are another two motions: feed motion vcz in the direction of the gear axis and axial shifting vcywhich is the movement of the honing wheel along its axis to refresh the working area. If the gear to be processed is a spur gear, the feed speed vcz could be a value independent with \u03c9g. But when to process a helical gear, there is a variation of \u03c9g which is linear to vcz. The variation could be expressed as \u0394\u03c9g = kgvcz. Then, \u03c9g \u00bc i \u03c6c; l\u00f0 \u00de\u03c9c \u00fe\u0394\u03c9g: For the axial shifting vcy, \u0394\u03c9c = kcvcz could also be determined"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000203_j.ijmecsci.2013.11.010-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000203_j.ijmecsci.2013.11.010-Figure7-1.png",
        "caption": "Fig. 7. Bending deformation of threads. (a) Cross section I-I shown in Fig. 6, (b) cantilever DO, (c) relationship among moment MO, shear force Q and axial force T for deformed shape.",
        "texts": [
            " In order to evaluate the items (1) and (2), a theoretical modeling for the bending deformation of threads should be needed. Also, as for the item (3), a theoretical evaluation for the bulging deformation of the rubber tube should be needed. Moreover, as for the item (4), a theoretical modeling for the relative sliding distance between threads and the tube should be needed. In the following section, three kinds of deformation mechanisms are introduced, and the evaluation equation for the relationship between the axial load F and the pressure difference P\u2032 is proposed. Fig. 7(a) shows a schematic of the vertical cross section I-I (see Fig. 6). As shown in Fig. 7(a), the force T acts on the thread during rotation, and the S-shaped bending deformation can be obtained between the intersections of threads. The portion where threads contact mutually is shown by the shaded portion in the figure. The width of this contacting portion is denoted by ac and is given by the following equation: ac \u00bc a= sin 2\u03b3: \u00f010\u00de It can be found from Eq. (10) that although the distance l0 between intersections remains unchanged during deformation, the width ac of the contacting portion increases as the angle \u03b3 deviates from 451",
            " 6(a), which acts between the adjacent threads, increases due to the tensile force T, and a larger frictional force must be overcome to maintain the rotation of the threads. Therefore, in order to take the effect of the rotation of threads into consideration, both the rotational resistances due to the frictional force among threads, and the change in the bending strain energy associated with this rotational behavior should be considered. Based on the symmetry of the thread deformation shown in Fig. 7(a), only half of the region DA, namely, the region DO is considered here, which can be analyzed by treating it as a cantilever of length xO \u00bc xA=2, as shown in Fig. 7(b). A force Q acts on the cantilever DO in the upward direction at the point E (xE \u00bc ac=2). The bending moment MO acting at the point O can be written in terms of the forces Q and T, and the bending moment MA as shown in Fig. 7(c) by Eq. (11): MO \u00bcMA Q \u00f0xF xO\u00de\u00feT\u00f0h\u00fe\u03b4\u00de=2; \u00f011\u00de where parameters h and \u03b4 represent the thickness of the thread and the gap among the threads, respectively. Based on the classical beam theory, the bending deflection curve for the cantilever is derived as follows. The moment M(x) along the length of the cantilever can be written by the following: M\u00f0x\u00de \u00bc MO\u00feT h\u00fe\u03b4 2 y\u00f0x\u00de Q \u00f0xO xE\u00de for 0rxrxE; MO\u00feT h\u00fe\u03b4 2 y\u00f0x\u00de Q \u00f0xO x\u00de for xErxrxO: 8>>< >>: \u00f012\u00de Therefore, the differential equations for the deflection y (in the following, denoted by yl and yr for two regions) of the cantilever are given as follows: EI d2yl dx2 Tyl \u00bc \u00bdMA\u00feT\u00f0h\u00fe\u03b4\u00de Q \u00f0xF xE\u00de \u00f013\u00de for 0rxrxE , and EI d2yr dx2 Tyr \u00bc Qx \u00bdMA\u00feT\u00f0h\u00fe\u03b4\u00de QxF \u00f014\u00de for xErxrxO",
            " When the point N is attached completely to the thread, and the thread deforms from the angle \u03b3 to \u03b3\u00fed\u03b3, the coordinate values termed xNt and yNt can be expressed as follows: xNt \u00bc k1xM\u00fek2s sin \u03b3; yNt \u00bc k2yM\u00fek1s cos \u03b3; \u00f030\u00de where parameters k1 and k2 are given by the following: k1 \u00bc jOD\u2032j jODj ffi1 sin \u03b3 cos \u03b3 d\u03b3; k2 \u00bc jOA\u2032j jOAj ffi1\u00fe cos \u03b3 sin \u03b3 d\u03b3: \u00f031\u00de On the other hand, when the point N is attached completely to the rubber, the coordinate values at the point N after deformation are denoted by \u00f0xNr ; yNr\u00de, and these values can be expressed as follows: xNr \u00bc k1\u00f0xM\u00fes sin \u03b3\u00de; yNr \u00bc k2\u00f0yM\u00fes cos \u03b3\u00de: \u00f032\u00de Therefore the magnitude of the relative sliding denoted by dutr can be written as follows: dutr \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0xNr xNt\u00de2\u00fe\u00f0yNr yNt\u00de2 q \u00bc 2s j cos 2\u03b3j sin 2\u03b3 d\u03b3: \u00f033\u00de It can be understood from Eq. (33) that there is no relative sliding between each thread and the rubber along the centerline of the thread, namely, s\u00bc0. Also, the magnitude of relative sliding is the same at two points which apart from the centerline with the same distance. Due to the S-shaped bending deformation, all threads would contact with each other at the junction points of their edges, for example, the points E and F shown in Fig. 7(a). Fig. 9 shows a contact region of the threads (described by the shaded portion). That is, for an intersection of threads, the threads contact each other at four junction points by each edge. Since the force applied to each junction point is Q=2, the moment Mf due to the frictional forces can be obtained as follows: Mf \u00bc \u03bcQa 2 1 sin \u03b3 \u00fe 1 cos \u03b3 ; \u00f034\u00de where \u03bc is the coefficient of friction between threads. Additional energies are required in order to produce bending deformation of the thread. The rotational resistant moment of a thread Md can be calculated from the following: Md \u00bc dUd 2d\u03b3 ; \u00f035\u00de where dUd denotes the increment of deformation energy, which is necessary for the deformation of four cantilevers surrounding the intersection point A, and is given by the following equation: dUd \u00bc 4 Z xO 0 EIy\u2033\u2202y\u2033 \u2202\u03b3 dx d\u03b3; \u00f036\u00de where y\u2033 is the second differential of the deflection obtained by the deflection of threads (Eqs",
            " Thus, the maximum contraction ratio \u03b7max for the thread having a three-circle cross section becomes larger than that having a rectangular cross section. (2) In this study, two kinds of braiding pattern of the threads are used as shown in Fig. 11. In Pattern 1, top and bottom threads alternate as they crisscross. On the contrary, in Pattern 2, one thread passes above (or below) two threads before they crisscross. The rotational resistance of the thread in Pattern 2 can be also analyzed by assuming a bending cantilever, as shown in Fig. 7(a). However, it should be mentioned that the braiding pattern affects the distance xA and xE shown in Fig. 7(a). For example, as for Pattern 1 model, xA \u00bc l0 and xE \u00bc ac=2, and as for Pattern 2 model, xA \u00bc 2l0 and xE \u00bc \u00f0l0=2\u00feac\u00de=2. Therefore, it can be understood that the rotational resistance for Pattern 2 is smaller than that for Pattern 1. Fig. 17 shows the comparison of \u03b7-P\u2032 curves for these two braided patterns. As can be seen from Fig. 17 that the maximum contraction ratio for Pattern 2 is larger than that for Pattern 1. However, Pattern 2 also has a demerit. The threads for Pattern 2 can easily slip because the contact force between threads is left\u2013right asymmetrical"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000651_1.3656586-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000651_1.3656586-Figure4-1.png",
        "caption": "Fig. 4 Film thickness in a spherical bearing",
        "texts": [
            " V Effect of External Pressurization Consider a hemispherical bearing with concentric surfaces, then H = 1. Let an orifice of infinitesimal opening be located a at the pole (<f> = 0). Due to symmetry, \u2014 = 0, then the steadycw state version of equation (31) is reduced to (43) d ( . , \u201e dP0 \\ \u201e The fluid film thickness depends on the position of the journal center relative to that of the bearing center. This relative position has two degrees of freedom; namely, an axial eccentricity ec and a radial eccentricity er as shown in Fig. 4. Thus h = C + e, cos 4> + e, sin <f> cos (8 \u2014 a); (37) a being the meridianal angular position of maximum film thickness. In the dimensionless form, one obtains H = 1 + tz cos 0 + tr sin 4> cos (8 \u2014 a) (38) IV Boundary Conditions Equation (31) is to be solved for a given set of initial and boundary conditions. Henceforth, only the steady-state problem of a plain hemispherical bearing will be considered; hence the initial condition will be disregarded. The 0-wise boundary condition is P{8, 4>) = P(8 + 2ir, 0 ) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003019_j.triboint.2017.04.018-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003019_j.triboint.2017.04.018-Figure5-1.png",
        "caption": "Fig. 5. Internal friction torque after 1 h and after 24 h.",
        "texts": [
            " However, it is interesting to notice that the internal friction torque of the RTB, lubricated with the candidate (B) oils (75W85-B and 75W90-B), were identical after 1 h and 24 h ( \u00b1 0.003%). In the case of the reference (A) formulations, the internal friction torque at the end of the tests is always lower than after 1 h, \u22127.6% for oil 75W90-A, \u22124.7% for oil 80W90-A and \u22125.8% for oil 75W140-A. This different behaviour of the candidate (B) and of the reference (A) axle gear oils, can be clearly observed in Fig. 5, where the internal friction torque is plotted for each axle gear oil formulation, after 1 h and at the end of the test. The surface topographies of the lower (stationary) raceways, of the new and used RTB, were measured at the end of the tests. The 3D roughness parameters of the surface topographies and the 2D roughness parameters of the roughness profiles extracted from those surface topographies are presented in detail in Appendix C. Table 5 presents the roughness parameter Rq (root mean square of the profile height) of the RTB raceways, as new and used (at the end of the tests)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003714_s12008-018-0520-6-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003714_s12008-018-0520-6-Figure1-1.png",
        "caption": "Fig. 1 Geometry and mesh of the model used in the present investigation",
        "texts": [
            " The interaction between laser power, laser speed, part geometry and how it affects thermal stress and final product properties is investigated. This approach enables design revision to the part geometry and selection of optimumAM process parameters. AnEulerian finite elementmodel was developed in this study using the commercial package, ANSYS, to analyze an additive manufacturing process subject to a moving laser heat beam approximated as Gaussian distribution. The geometry of the model used is depicted in Fig. 1. The mesh consisted of 8-node element with 29,852 elements. The ANSYS program uses the Newmark time integration method to solve the governing heat conduction equations at discrete time-points. The present model was built by superimposing 4 single layers, each of 1 mm thick and total height of 4 mm. This object was constructed on a solid substrate with base dimensions of 10mm\u00d710mm and thickness of 4mm. Themoving laser heat source was approximated using Gaussian volumetric distribution function as follows [10]: S C2 \u00d7 exp \u2212{ (x \u2212 vx \u00d7 t)2 + (y \u2212 vy \u00d7 t)2 } C2 1 (1) where S is the heat flux on the surface, C1 is the laser beam radius (1 mm in this study), C2 is the source power intensity (= 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002435_iros.2015.7353998-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002435_iros.2015.7353998-Figure3-1.png",
        "caption": "Fig. 3. An omni-directional wheel represents the bicycle\u2019s back wheel",
        "texts": [
            " The effect of tissue deformation can be interpreted as letting the needle have sideways movements orthogonal to the insertion direction. These movements are modeled by slippage of the bicycle wheels. As explained in the previous section, the original kinematic bicycle equations were derived by imposing the pure roll and non-slip constraints on the wheels. In order to allow for sideways motion of the needle, we replace the back wheel of the bicycle with a wheel that is able to move in two directions. Omni-directional wheels shown in Figure 3 can move independently in two orthogonal directions. Such wheels satisfy the wheel plane constraint (3) but violate the non-slip constraint (4). Orthogonal to the wheel plane, the motion constraint will be Bvy \u2212 \u03c9R = 0 (11) in which \u03c9R denotes the rotation velocity of the rollers. This is the degree of freedom added to the system allowing the wheel to have lateral movements. Here, we replace the back wheel of the bicycle with the omni-directional wheel of Figure 3, as shown in Figure 4. In this Figure, \u03b2, ` and \u03b1 denote the fixed front wheel angle, the distance between the two wheels and the rotation angle of the needle tip in body frame {B}, respectively. This rotation of the needle tip from the insertion direction by angle \u03b1 is due to the lateral movement of the back wheel causing the needle to be tangent to the non-circular path. The inputs u1 and u2 denote the insertion velocity along the z axis of frame {B} (which equals Bvz) and the rotation velocity of the needle about its axis, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000683_icra.2013.6631258-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000683_icra.2013.6631258-Figure5-1.png",
        "caption": "Fig. 5. 4-link manipulator.",
        "texts": [
            " Then, in order to consider dynamic reconfiguration measure of the whole manipulator-links, we define a index named dynamic reconfiguration manipulability shape index (DRMSI) as follows: 1WDR = n\u22121\u2211 j=1 aj 1VDRj (24) Here, aj is unit adjustment between different dimension. In this paper, singular-values increase a hundredfold to enlarge value of ellipsoid, compared to ellipse or line segment. In this section, we introduce a numerical example of the proposed DRM for posture of a 4-link manipulator in Fig. 5. Head Upper body Upper arm Lower body Hand Lower arm Waist Upper leg Lower leg Middle body Foot : Number of link : Number of joint Fig. 8. Definition of humanoid\u2019s link, joint and angle number. TABLE I PHYSICAL PARAMETERS Link li mi di Head 0.24 4.5 0.5 Upper body 0.41 21.5 10.0 Middle body 0.1 2.0 10.0 Lower body 0.1 2.0 10.0 Upper arm 0.31 2.3 0.03 Lower arm 0.24 1.4 1.0 Hand 0.18 0.4 2.0 Waist 0.27 2.0 10.0 Upper leg 0.38 7.3 10.0 Lower leg 0.40 3.4 10.0 Foot 0.07 1.1 10.0 Total 1.7 63.8 Length, mass and coefficient of viscous friction of each link and joint are set to be 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001466_1350650118778655-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001466_1350650118778655-Figure10-1.png",
        "caption": "Figure 10. Comparison of predicted mean coefficient of friction, numerical simulations and experimental results at 8.3 m/s and 183.4 Nm for Cmod gears.",
        "texts": [
            "10 has been obtained empirically from twin disc measurements while the reference shear stress used in the current paper has been obtained analytically in Bair and Winer.24 The latter departs from a the limiting stress pressure behaviour of the type L \u00bc p, which only includes the influence of pressure (at constant temperature), as verified experimentally in high pressure shear stress measurements, while Ziegltrum et al. also include the influence rolling velocity and slip ratio. In order to give further insight into the proposed Eyring stress model, E \u00bc 2 = , the influence of the limiting stress pressure coefficient, , has been analysed in Figure 10. The latter has been varied 3 10 3 from its reference value, which is approximately the fluctuation of this coefficient at the range of operating temperatures in the contact. As it can be seen, the Eyring stress, and therefore the mean friction coefficient, is ruled primarily by the piezoviscosity coefficient, , as the influence of varying is negligible. In addition, Figure 10 has been completed with the experimental measurements from Hinterstoi\u00dfer40 for the same gear set, operating conditions and oil types. The experimental results show that mean friction coefficient is higher in all cases, suggesting that mixed lubrication regime is prevailing along the line of action. This affirmation is supported by the fact that both the numerical and the analytical models predict film thicknesses just below 0.2mm which is the value of the actual surface roughness in Hinterstoi\u00dfer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002426_j.proeng.2015.12.025-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002426_j.proeng.2015.12.025-Figure2-1.png",
        "caption": "Fig. 2. Power factors of friction in contact of a wheel to a ground.",
        "texts": [
            " Let's consider it more in detail. Each elementary force of friction dP is opposite to a direction of relative sliding. Each point of contact has its direction of speed sliding. It allows to reduce a problem in formation of forces in contact to kinematics of relative sliding. How are forces of friction in contact of a wheel and road on turn formed? We shall consider it on the basis of the mathematical theory of friction [10]. The local system of coordinates Y is connected with the center of contact (figure 2). The axis is directed along movement of a wheel, but the axis Y is perpendicular. The elementary point with the area dF has in local system of coordinate , . Its speed of sliding V is directed perpendicularly to distance r up to the instant center of speeds (point C with coordinates x, y). Projections of elementary force of friction dP to the axes Y are equal: sinxdP dP and cosydP dP . (2) We shall consider Fig. 1. Sum elementary forces of friction in contact: (a) elementary forces dPpi; (b) elementary forces dPmi; (c) elementary forces dPi"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001585_aim.2018.8452392-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001585_aim.2018.8452392-Figure14-1.png",
        "caption": "Figure 14. Joint part of previous robot",
        "texts": [
            " Compared with the previous unit part, the diameter of the artificial muscle is thinned by reducing the size of the artificial muscle fastener (Fig.11). As a result, it was confirmed that the contraction inside the tube improved by 70%. The contraction characteristics of the existing robot unit are shown in Fig.12, and the shrinkage characteristics of the high contraction unit of the proposed robot are shown in Fig.13. The proposed robot also has a joint structure that imitates a universal mechanism (Fig.14). A spring is also contained inside the join in order to increase the rigidity. As shown in Fig.15, the robot was shortened, and the density per length of the joint part was increased. Therefore, even if the joint part is short, it can pass through the elbow pipe. The robot intakes and discharges air from the unit by opening and closing the solenoid valve. The solenoid valve is opened and closed by a microcomputer. The solenoid valves are arranged in the tail and arranged on the cable. Since a certain width is generated in the size of the solenoid valve, it is possible to select one having an appropriate effective sectional area"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000748_tie.2013.2276025-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000748_tie.2013.2276025-Figure2-1.png",
        "caption": "Fig. 2. Complete view of the device for induction heating of rotating billets.",
        "texts": [
            " This paper represents a natural continuation of previous works [19] of the authors where they described modeling and experimental verification of induction heating of a thin-wall aluminum pipe. It contains a complete mathematical model of the process, its numerical solution carried out by own code based on a fully adaptive higher order finite-element method, evaluation of the principal operation characteristics, and their experimental verification. The complete view of the proposed experimental device is shown in Fig. 2. It consists of an induction motor driving the billet and an active heating part consisting of a magnetic circuit with fixed high-parameter permanent magnets. The cross section of this active part, together with its principal dimensions in millimeter, is shown in Fig. 3. The indicated orientation of particular permanent magnets was proposed after a series of preliminary computations aimed at the evaluation of operation parameters and efficiency of the device. This paper deals with numerical modeling of the principal characteristics of the device, namely, the time evolution of the temperature in the billet, balance of power, efficiency, and their experimental verification"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003981_j.jfranklin.2018.10.041-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003981_j.jfranklin.2018.10.041-Figure1-1.png",
        "caption": "Fig. 1. The 2DOF and 3DOF planar robots in MHRS.",
        "texts": [
            " Clearly, one can also obtain that based on lemma 2 such that one can say that the developed controllers can also be employed to address control problems under cyclic graphs. Thus, they can be used to enhance the feasibility and universality of time-delay MHRS in practical engineering applications. 4. Simulations To demonstrate the validity of the proposed controllers, we consider the MHRS with a master robot 0 and six different degrees-of-freedom (DOF) planar robots, namely, three 2DOF and three 3DOF slave robots. The mechanical structures are shown in Fig. 1 . The physical parameters of the slave robots are shown in Table 2 , where m i , i \u2208 { 1 , . . . , 6 } , denotes the mass, l i denotes the length, r i denotes the distance between the joint and the center of mass, I i is the moment of inertia around its center of mass. The kinematic and dynamic c 2 M w M C g M w M onstant parameters of the MHRS are shown in Table 3 [31,37] . Thus for 2DOF robots i \u2208 {1, , 3}, i ( q i ) = [ M i11 M i12 M i21 M i22 ] , C i ( q i , \u02d9 qi ) = [ C i11 C i12 C i21 C i22 ] , g i ( q i ) = [ g i1 g i2 ] , here i11 = \u03d1 di1 + 2 \u03d1 di2 cos q i2 , M i12 = M i21 = \u03d1 di3 + \u03d1 di2 cos q i2 , M i22 = \u03d1 di3 , i11 = \u2212\u03d1 di2 \u0307 qi2 sin q i2 , C i12 = \u2212\u03d1 di2 ( \u0307 qi1 + \u02d9 qi2 ) sin q i2 , C i21 = \u03d1 di2 \u0307 qi1 sin q i2 , C i22 = 0, i1 = \u03d1 di4 g cos q i1 + \u03d1 di5 g cos ( q i1 + q i2 ) , g i2 = \u03d1 di5 g cos ( q i1 + q i2 ) , g = 9 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003464_s12555-017-0280-2-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003464_s12555-017-0280-2-Figure1-1.png",
        "caption": "Fig. 1. Measurement of relative orientation angle.",
        "texts": [
            " Problem statement for single-integrators Consider the following N agents on the plane: p\u0307i = vi, \u03b8\u0307i = \u03c9i, (2) where pi \u2208 R2 and vi \u2208 R2 are the position and the linear velocity input of agent i with respect to the global reference frame, \u03b8i \u2208 R is the orientation angle of the agent relative to the global reference frame, and \u03c9i \u2208 R is the angular velocity input for i = 1, . . . ,N. Let g \u2211 denote the global reference frame. Further let i \u2211 denote the local body-fixed reference frame of agent i that is located at pi with its orientation angle \u03b8i relative to g \u2211 (see Fig. 1). It is assumed that the interaction among agents agents can be modeled by a graph G = (V,E ,A), the nodes and the edges of which denote the agents and the interaction among them, respectively. Let G denote the interaction graph of the agents. Assume that agent i measures the relative positions of its neighbors with respect to i \u2211. By adopting a notation in which superscripts are used to denote reference frames, agent i measures the following relative positions: pi ji := pi j \u2212 pi i, j \u2208Ni, (3) where pi i and pi j denote the positions of agents i and j with respect to i \u2211. Define PV : R\u2192 (\u2212\u03c0,\u03c0] as PV(x) := 2k\u03c0 +x, where k is an integer such that 2k\u03c0 +x \u2208 (\u2212\u03c0,\u03c0]. Assume that the relative orientation angles, \u03b8 ji := PV(\u03b8 j \u2212\u03b8i) , j \u2208Ni, (4) are available to agent i. The relative orientation angles can be calculated by agent i, for instance, as follows. As depicted in Fig. 1, agent j measures \u03b4i j and transmits the measured value to agent i by communication. Agent i measures \u03b4 ji and calculates the relative orientations of its neighbors based on that PV(\u03b8 j \u2212\u03b8i) \u2261 PV(\u03b4 ji \u2212\u03b4i j +\u03c0). Note that G describes which relative information is used by whom, rather than the sensing topology itself. In below, a angular velocity input will be designed to drive the orientation angles to a common constant value. Let \u03b8\u221e \u2208 R be the limiting value. Let c \u2211 denote the reference frame located at the origin of g \u2211 with its orientation angle \u03b8\u221e relative to g \u2211 (see Fig. 1). Denote by pc i the position of agent i with respect to c \u2211. Assume that agent i estimates pc i . Denote by p\u0302c i the estimated value of pc i . The following relative variables are available to agent i: p\u0302c ji := p\u0302c j \u2212 p\u0302c i , j \u2208Ni. (5) Note that the availability of measurements (3), (4), and (5) is related to the assumption on the sensing and communication capability of the agents (2). The formation control problem for (2) is then stated as follows: Given p\u2217 = [p\u2217T 1 \u00b7 \u00b7 \u00b7 p\u2217T N ]T \u2208 R2N , design vi, \u03c9i, and a position estimation law to achieve \u03b8 j \u2212\u03b8i \u2192 0, p\u0302c j \u2212 p\u0302c i \u2192 pc j \u2212 pc i and pc j \u2212 pc i \u2192 p\u2217j \u2212 p\u2217i by using the measurements (3), (4), and (5) for i = 1, ",
            " Since it is assumed that maxi\u2208V \u03b8i(t0)\u2212mini\u2208V \u03b8i(t0)< \u03c0 in this paper, the orientations of the agents need to be roughly aligned at the initial instance, which is not desirable. However, this condition is not so restricted in practice because engineered systems often require some initial setup. As discussed above, there exists \u03b8\u221e \u2208 R such that \u03b8(t) exponentially converges to \u03b8\u221e1N as t \u2192 \u221e under the given condition. Based on the existence of \u03b8\u221e, consider the reference frame located at the the origin of g \u2211 with its orientation angle \u03b8\u221e relative to g \u2211 (see Fig. 1). Denote the reference frame by c \u2211. Note that each of the local bodyfixed frames of the agents exponentially aligns to c \u2211. This orientation alignment allows the agents to have a common sense of orientation. Based on the obtained common sense of orientation, a position estimation based formation control law can be proposed for the agents (2). Let vi i denote the linear velocity input of agent i with respect to i \u2211. To allow the agents to estimate their positions with respect to c \u2211, consider the following position estimation law: \u02d9\u0302pc i = kp\u0302 \u2211 j\u2208Ni ai j [ (p\u0302c j \u2212 p\u0302c i )\u2212 pi ji ] + vi i, (8) where kp\u0302 > 0",
            " It should be noted here that k\u03b8 should be large enough so that the dynamics of (7) is faster than that of (8). Further, kp\u0302 and kp should be designed such that the dynamics of (8) is faster that that of (9). Note that (7), (8), and (9) can be implemented by using the measurements (3), (4), and (5) in the local body-fixed reference frames. For the purpose of the stability analysis, describe the dynamics of the agents (2) in the reference frame c \u2211. Define e\u03b8i := \u03b8\u221e \u2212\u03b8i. Since i \u2211 has its orientation angle \u2212e\u03b8i relative to c \u2211 (see Fig. 1), it is obvious that vc i = R\u22121(e\u03b8i)v i i, pi ji = R(e\u03b8i)(pc j \u2212 pc i ), where R(e\u03b8i) \u2208 R2\u00d72 denotes the rotation matrix by e\u03b8i . Let e\u03b8 = [e\u03b81 \u00b7 \u00b7 \u00b7e\u03b8N ] T and \u0393(e\u03b8 ) := diag(R(e\u03b81), . . . ,R(e\u03b8N )). Further let pc = [pcT 1 \u00b7 \u00b7 \u00b7 pcT N ]T and p\u0302c = [p\u0302cT 1 \u00b7 \u00b7 \u00b7 p\u0302cT N ]T . To obtain the error dynamics, let epc := p\u2217 \u2212 pc, p\u0303c := pc \u2212 p\u0302c, and es := [eT pc p\u0303cT ]T . The following error dynamics is obtained: e\u0307s = Ases +\u2206As(e\u03b8 )es +Ds(e\u03b8 ), (10a) e\u0307\u03b8 =\u2212k\u03b8 Le\u03b8 , (10b) where As := [ \u2212kpI2N \u2212kpI2N 0 \u2212kp\u0302(L\u2297 I2) ] , \u2206As(e\u03b8 ) := [ I2N \u2212\u0393\u22121(e\u03b8 ) ][ kpI2N kpI2N \u2212kpI2N \u2212kpI2N ] \u2212 [I2N \u2212\u0393(e\u03b8 )] [ 0 0 k p\u0302(L\u2297 I2) 0 ] , and Ds(e\u03b8 ) := [ I2N \u2212\u0393\u22121(e\u03b8 ) ][ 0 k p\u0302(L\u2297 I2)p\u2217 ] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003467_s40997-018-0184-7-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003467_s40997-018-0184-7-Figure6-1.png",
        "caption": "Fig. 6 Dynamic model of a helical gear pair",
        "texts": [
            " LTCA is a method for simulating the contact process of gears under load conditions, which provides the data for gear dynamics analysis, especially for the calculation of inner excitations (Wang et al. 2017; Wang and Shi 2017). The model of LTCA for a helical gear pair is shown in Fig. 5. The mathematical description is shown in Eq. (10). The detailed solving process of LTCA can be found in Conry (1971, 1973). Wang et al. (2017) describes in some detail the process of calculating internal excitations by the data obtained by LTCA. Min P2n\u00fe1 j\u00bc1 Xj \u00bdF \u00bdp \u00fe \u00bdZ \u00fe \u00bdd \u00fe \u00bdX \u00bc \u00bdw \u00bde T \u00bdp \u00fe X2n\u00fe1 \u00bc P S:t pj ; dj ; Z ; Xj 0 pj \u00bc 0 or dj \u00bc 0 8>>>< >>>>: \u00f010\u00de Figure 6 shows an 8-degree bending-torsion-shaft coupled vibrating system for a helical gear pair, the eight degrees are described as df g \u00bc x1y1z1h1x2y2z2h2f gT \u00f011\u00de According to Newton\u2019s second law, the equations of vibrating system are described as m1\u20acx1 \u00fe c1x _x1 \u00fe k1xf1x\u00f0x1\u00de \u00bc klF1y \u00f012\u00de m1\u20acy1 \u00fe c1y _y1 \u00fe k1yf1y\u00f0y1\u00de \u00bc F1y \u00f013\u00de m1\u20acz1 \u00fe c1z _z1 \u00fe k1zf1\u00f0z1\u00de \u00bc Fz \u00f014\u00de I1\u20ach1 \u00fe F1yr1 s1klF1y \u00bc T1 \u00f015\u00de m2\u20acx2 \u00fe c2x _x2 \u00fe k2xf2x\u00f0x2\u00de \u00bc klF1y \u00f016\u00de m2\u20acy2 \u00fe c2y _y2 \u00fe k2yf2y\u00f0y2\u00de \u00bc F1y \u00f017\u00de m2\u20acz2 \u00fe c2z _z2 \u00fe k2zf2z\u00f0z2\u00de \u00bc Fz \u00f018\u00de I2\u20ach2 F2yr2 \u00fe s2klF1y \u00bc T2 \u00f019\u00de The detailed solving process of the dynamic model can be found in Wang and Shi (2017)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002867_arso.2016.7736297-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002867_arso.2016.7736297-Figure2-1.png",
        "caption": "Figure 2. The main processing methods of the blade",
        "texts": [
            " 3) Completed path planning after calculating the posture of the processing points based on the domain information of processing points. Normal vector at any point on the surface could be described as the following formula u v u v r r N r r    (3) Wherein u rr u   and v rr v   were called vector coordinates, respectively as the tangent vector of lines u and lines v of the curved surface on processing points. For the processing of the blade, according to the processing methods that could be divided into the horizontal line cutting method and vertical line cutting method, which were shown in the Fig.2. Considering the characteristics of the two methods, this paper took the horizontal line cutting method to process the blade along cross-sectional profile under the existing conditions. The horizontal line cutting method needed to ensure the path number according to the processing line spacing, that was to calculate how much of the processing paths for the entire workpiece being machined. As shown in Figure 3, the number of the machining paths could be obtained by the following formula: L P d         (4) Where P was the determined number of the processing path, L was the length of the entire workpiece-surface, d was the bonded width between the polishing wheel and the workpiece, \u03b4 was the overlap region width between two adjacent paths"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002913_i2016-16121-7-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002913_i2016-16121-7-Figure3-1.png",
        "caption": "Fig. 3. Winding of the dowser texture by a rotating magnetic field.",
        "texts": [
            " Knowing that d is orthogonal to the meniscus (homeotropic boundary conditions at the air/nematic interface) it must thus have the orthoradial orientation in the vicinity of the hedgehog which means that the hedgehog has rather the \u201ccircular\u201d structure represented in fig. 2 characterized by the phase \u03c80 = \u2212\u03c0/2. Inspired by previous experiments with umbilics [21], we submitted the dowser state with the residual hedgehog to a rotating magnetic field with the aim to wind the phase \u03c80 of the hedgehog. Results of a typical experiment are shown in fig. 3. In the initial state (fig. 3a) the four isogyres are almost straight except in the vicinity of the hedgehog where they have spiral shapes. This means that the dowser field is radial (\u03c8 = 0) except in the vicinity of the hedgehog where its phase \u03c8(r) varies between 0 and \u2248 \u2212\u03c0/2. In this experiment, the sample is thick (h = 0.1mm) and we are observing the middle part of the sample, where the thickness gradients are low, so the effect of cuneitropism \u2014alignment of the dowser field along the thickness gradient, as described in ref. [15], is less pronounced and compression of isogyres into 2\u03c0 walls is not seen. Application of a horizontal magnetic field B (see fig. 3b) to the radial dowser field results in formation of two \u03c0-walls connected to the hedgehog. The magnetic field starts then to rotate in the counterclockwise direction slowly enough to drag the dowser field with it. In figs. 3c-f we show the result of incremental rotation of B. Each \u03c0 turn of the dowser field in the sample centre driven by the magnetic field adds \u03c0 to the phase of the hedgehog \u03c80. At the same time the phase of the dowser field \u03c8(R) = 0 remains fixed by the homeotropic anchoring at the meniscus"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001061_icieect.2017.7916559-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001061_icieect.2017.7916559-Figure2-1.png",
        "caption": "Fig. 2, Propellers",
        "texts": [],
        "surrounding_texts": [
            "The standout amongst the most imperative piece of Quad Copter is its frame that supports engines and different hardware and keeps them from the Vibrations. Arm: Arm can be produced using any material like CF, PVC pipes, Aluminum or Wood, however, it has enough quality to withstand effect and harsh landings. Casing weight ought to be around 200-250 grams. Center Plate: Center plate holds the arms and backings Flight controller, collector and different sensors. We can utilize 2-3 mm glass fiber, plywood, aluminum sheet or any plastic sheet material yet verify they are firm, lightweight. Size: There is no standard size of the frame yet for medium size quad copter 450 mm to 550 mm engine to engine edge is sufficient. The frame specifications are given below. B. A1otor/~ngine One of the critical parts of Quad-Copter is its motor. It's a part of a power system. In fact entire power system is based on the choice of motor. To lift 1000 grams copter there will be a need of aggregate 2000 grams of thrust. It has 4 motor, so every motor ought to have the capacity to create at least 500grams of thrust to fulfill the need. 4 motors x 500grams push = 2000grams push. We utilize Brushless motor model MX2212. It has following features . 600g max thrust / motor which gives 2400g total thrust to lift lOA of maximum current which is easy to supply from battery"
        ]
    },
    {
        "image_filename": "designv11_13_0003418_02670836.2018.1466944-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003418_02670836.2018.1466944-Figure2-1.png",
        "caption": "Figure 2. Model of geometry and sample position.",
        "texts": [
            " The subscripts s and v denote the nucleation at themould surface and in the bulk of the liquid, respectively (Figure 1). Two functions were used to describe heterogeneous nucleation at the mould surface and in the bulk of the liquid [21]. After nucleation occurs, the dendrite growth in the casting is dependent on the total undercooling of the dendrite tip. The relationship between the growth rate of dendrites v( T) and undercooling T, is expressed by the following equation [26]: v( T) = 1.64 \u00d7 10\u22126 T2 + 1.03 \u00d7 10\u22127 T3 (9) Finite volumemesh and boundaries conditions Figure 2 shows the finite volume mesh for the 3-D geometry model and sample positions. For the symmetry of the geometry, half the casting (625mm\u00d7 210mm\u00d7 8000mm) was presented in the simulation. The CAmethod applies the non-steady state algorithm. Therefore, the Lagrangian algorithm could be used for both FE and CA calculations. Using the boundary movement method to calculate the distribution of the solidification temperature, coupled with the CA method to calculate grain formation, the cooling conditions between the crystalliser and casting should be translated along the lateral surface of the nonmoving slab ingot [15]. The initial position of the crystalliser, which is at the bottom of the ingot shown in Figure 2, moves up at a pulling speed that is set at 2.35\u00d7 10\u22124 m s\u22121 in this simulation. The calculation zone automatically generated a regular network of square cells at 200 \u03bcm spacing, to simulate the formation of the macroscopic grains using different bulk nucleation parameters ( Tv,max, Tv,\u03c3 , nv,max). The bulk nucleation parameters were determined by comparing the experimental results of different crosssections, which were 100, 200, 600, 800 and 1000mm distance from the bottom of the ingot, respectively",
            " The heterogeneous nucleation parameters at the surface, which only determined the nucleation at the mould wall, were kept constant for the sake of simplicity. The values of nucleation parameters at the surface were ( Ts,max = 0.5 K, Tv,\u03c3 = 0.1 K, ns,max = 2\u00d7 106 m\u22122) [16,30]. In this study, the simulation used different bulk nucleation parameters in which the cross-section was 800mm from the bottom of the slab ingot. The simulated cross-sections were compared to the micrographic samples and were in different positions (Figure 2). To determine the bulk nucleation parameters of the simulated macroscopic grains of the large-scale titanium slab ingot, the effects of bulk nucleation parameters need to be further studies, as described in a previous study [26]. Effect of the differentmean nucleation undercooling onmacroscopic grains Figure 4 shows the calculation of the formation of the macroscopic grains with different solidification times (t = 1160, 1380, 1580 s) using four values of mean nucleation undercooling ( Tv,max = 20, 15, 10, 5 K), Figure 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002173_detc2014-35099-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002173_detc2014-35099-Figure13-1.png",
        "caption": "Figure 13 Tessellated CAD model approximation of volumetric error with adaptive slicing",
        "texts": [
            " The step (ds) for an adaptive layer generation was set as 0.001mm. Figure 11 demonstrates the effect of the developed volumetric error control adaptive slicing procedure on the model relative to uniform layers. The change in volumetric error shown in the Fig.12 appears to be due to shallow slope surfaces being divided into the layers of variable volumetric error by the changing of the subsequent layer thickness. The implementation of \u201eVolumetric Error Control through Adaptive Slicing\u201f can be seen more clearly as described in Fig. 13. It shows the tessellated CAD model approximation of a volumetric error of an adaptively sliced symmetric model. The role of layer thickness variations to control the volumetric error using adaptive slicing of tessellated CAD model is presented in Figs. 14 & 15. Figure 14 is presented with layer thickness as a function of part height. Figure 15 represents 7 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use the variation of volumetric error for the particular layer with the part height"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002889_s40436-016-0158-1-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002889_s40436-016-0158-1-Figure2-1.png",
        "caption": "Fig. 2 The three-dimensional model of the spindle unit",
        "texts": [
            " In this case, the front motorized spindle bearing is fixed, and the rear bearings can move in the axial direction. The boundary conditions are applied to the front bearing shaft shoulder by which the front bearings are elastic supported, and the stiffness of the elastic support is the axial stiffness of the front bearings. In addition, it is considered that the material of the front and rear balancing rings are the same as the spindle, and the simplified method of the spindle unit is not different as the thermal-structure coupling model. The simplified model is shown in Fig. 2. To obtain the thermal characteristic of the simplified model, the ANSYS Workbench 15.0 commercial code is applied. The environment temperature of the simulation is 24 C and the rotating speed is 20 000 r/min. The internal temperature of the motorized spindle is shown in Fig. 3. As can be seen from Fig. 3, the main heat sources of the motorized spindle are the rotor, the stator, and the bearings. The front cooling water passage takes away a portion of the heat thus the surrounding temperature is relatively low"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003116_s11465-017-0452-z-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003116_s11465-017-0452-z-Figure9-1.png",
        "caption": "Fig. 9 Sketch map of the transmission chain of the investigated WT",
        "texts": [
            "1 Structure of the investigated WT In wind power generation, wind power is used to drive the windmill blades, and the rotational speed is increased by the speed-increase gearbox to actuate the generator and consequently produce electricity. The planetary gearbox is the core component for unit movement and energy transfer. However, gearbox fault is widely recognized as the leading issue in the condition monitoring of WT transmission systems [63]. A structural sketch of the investigated planetary gearbox of a typical 1500 kW WT is shown in Fig. 9. The transmission chain of the WT consists of two-stage planetary gears and one-stage fixed-shaft gears. The transmission of the input torque to the generator is performed by the gearboxes, which transmit the torque from the windmill blades to the planet carrier of the planetary gears and fixed-shaft gears. Table 1 lists the structural parameters of the rotating components in the transmission chain. The transmission ratios and orders of the transmission chain at each stage can be calculated using the structural parameters presented in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000196_j.jbiomech.2012.11.025-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000196_j.jbiomech.2012.11.025-Figure2-1.png",
        "caption": "Fig. 2. A cartoon depiction of a sperm in (A) where the tail is the flagellum and the circular structure is the cell body or head. The internal structure of the flagellum is called the axoneme and a cross section in the plane is depicted in (B). The axoneme consists of nine sets of microtubule doublets around the circular structure and a central pair. Dynein arms reach to attach to the next microtubule doublet creating crossbridges. The simplified flagellar geometry used in the model, the flagellar centerline is depicted in (C).",
        "texts": [
            " Since this is a fluid\u2013structure interaction model, the kinematics and trajectories that are achieved will depend on the surrounding viscous fluid and the properties of the elastic filament. We investigate the sensitivity of sperm trajectories to the frequency of Ca2\u00fe oscillations, accounting for the relevant hydrodynamics. In this coupled model, we observe that longer periods of Ca2\u00fe oscillation corresponds to circular paths with greater drift. In contrast, shorter periods of Ca2\u00fe oscillations corresponded to tighter search patterns. The interior of the sperm flagellum is a circular structure consisting of nine microtubule doublets surrounding a central pair (Fig. 2(B)). The sperm geometry, depicted in Fig. 2, will be simplified by modeling an elastic flagellum denoted by X\u00f0r,t\u00de \u00bc \u00f0x\u00f0r,t\u00de,z\u00f0r,t\u00de\u00de, corresponding to the centerline, where r is a parameter initialized as arc length and t is time. We simplify the flagellum by only capturing the centerline, corresponding to a 1-D filament that is restricted to swimming in the xz plane. This corresponds to planar flagellar bending, which can be assumed for sea urchin sperm (Guerrero et al., 2011). We note that we are not accounting for the cell body, which will cause small changes in fluid velocity, swimming speed, and trajectories as shown previously in experiments and computational models (Brokaw, 1965; Gillies et al",
            " When the intracellular Ca2\u00fe increases, an asymmetrical waveform has been observed, corresponding to an increased angle in the principle bend direction and a decreased angle in the reverse bend direction (Brokaw, 1979). At high Ca2\u00fe , this asymmetrical waveform is no longer truly sinusoidal. In this model, we are able to account for this asymmetric bending as the Ca2\u00fe increases by choosing kb in Eq. (2b) based on the sign of the preferred curvature. When kb,1 4kb,2, this defines the principle bending direction as the one with positive curvature. The actual location of where Ca2\u00fe is acting within the axoneme to modify bending is not completely known. Dyneins (Fig. 2(B)), chemo-mechanical ATPases or force generators, cause sliding of the internal microtubule doublets, which is converted into bending (Lindemann and Lesich, 2010; Woolley, 2010). It is hypothesized that Ca2\u00fe acts by binding to proteins such as calmodulin, which interact directly or indirectly with dynein complexes, to cause a conformational change in the dyneins that results in an asymmetric change in dynein force generation, causing asymmetrical flagellar bending (Lindemann, 2007; Smith, 2002; Smith and Yang, 2004; Yang et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002210_1.4029828-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002210_1.4029828-Figure11-1.png",
        "caption": "Fig. 11 Models of nonuniform and nonhomogeneous rotating ring with: (a) local imperfection, (b) material periodicity, (c) geometric periodicity, and (d) elastic foundation periodicity",
        "texts": [
            ", the system becomes stiffer at low frequencies. Figures 9(c) and 10(c) show that the structural damping attenuates the amplitudes of resonance peaks without modifying resonant frequencies. As expected, the amplitude of the rigid body mode peak is not influenced by the internal damping of the ring. 3.3 Rotating Ring With Imperfections. One of the objectives of this work is to propose a method able to simulate rotating rings with nonuniformities and nonhomogeneities. Some examples are illustrated in Fig. 11. Imperfections may be produced by differences in material or geometric properties within a structure or by boundary conditions. Internal pressure and the structural damping are not considered in this section. The structure is excited in the radial direction and the radial displacement solution is computed at a node diametrically opposed to the excited one. In this section, when specified, uniform refers to the cross section dimensions bU\u00bc 0.15 m and hU\u00bc 0.002 m, listed in Table 3, and homogeneous to material with averaged density and Young\u2019s modulus: qH\u00bc (qI\u00fe qII)/2 and EH\u00bc 2/(1/EI\u00fe 1/EII), respectively, where indices I, II, and H correspond to materials listed in Table 4. First, the effect of a local imperfection was investigated (Fig. 11(a)). A uniform ring of radius R\u00bc 0.20 m was constructed with two elements: one with an angular length of 350 deg and made of material I, and another one with angular length of 10 deg and made of material II. The dynamic response of this ring is compared with that of a uniform ring made of material I. In Fig. 12, the forced responses of the rings with and without imperfection are compared. From 0 to 100 Hz, the differences are not clear. However, at higher frequencies, the flexural modes of the ring with imperfection become nondegenerated, i",
            " The full structure is composed of 24 elements of equal angular length (15 deg) and alternating properties. Differences in the material, cross section geometry, and elastic foundation are investigated. The rotational speed was kept constant at 50 rad/s. In the case, the periodicity is in the material properties, materials I and II (from Table 4) are used and the geometry of the ring is uniform. When the ring presents a periodic nonuniformity, homogeneous elements with cross sections 1 and 2 (Table 3) are used (Fig. 11(c)). In the last case investigated, uniform homogenous elements supported or not on an elastic foundation are alternated (Fig. 11(d)). An equivalent uniform and homogeneous ring without elastic foundation is used as baseline. Each periodicity produces different band gaps (frequency bands where no propagating waves exist), with different depths (attenuation levels of the forced response) and widths (of the frequency bands), which can be clearly observed when compared with the baseline, i.e., the equivalent uniform homogeneous ring (Fig. 13). Material and geometric periodicities create stop bands in all frequency ranges, and their depth and width become larger as frequency increases, while the elastic foundation periodicity creates band gaps at low frequencies only, with larger depth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002199_1.4892628-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002199_1.4892628-Figure5-1.png",
        "caption": "FIG. 5. (a) shows the schematic of a real gecko seta. The seta is cantilevered from its root and bends under the actions of bending forces (Ft and Fn give rise to bending force FtL sin\u00f0h\u00de FnL cos\u00f0h\u00de) and moments (Mc) applied at its tip. While adhered to a surface, the contact pad must be rotated to remain parallel to the surface. (b) shows the beam like model of a seta from (a) mapped to a system of two torsional springs (kc and kg). In both (a) and (b), the left hand figure shows the undeformed seta and the right hand side the seta supporting a load.",
        "texts": [
            " The heuristic model is overly simplistic and includes no linear extension of the seta. This means that it would require no work to detach the seta if the force is applied along ho. To correct this, we consider a second model that is more representative of a gecko seta and that permits us to examine work of de-adhesion. This model has two significant differences, first it assumes that the seta is extensible, and second that it supports a moment both at its tip, and at its root where it joins the gecko\u2019s foot, as shown in Fig. 5(a). We impose the boundary condition that as the seta flexes the contact pad must remain horizontal to stay bonded to the surface. This gives the condition that af\u00fe am\u00bc 0, where af, and am are the rotation angle of the seta tip due to bending force, \u00f0FtL sin\u00f0h\u00de FnL cos\u00f0h\u00de\u00de, and a moment Mc, respectively. We consider the seta as a curved beam with a varying moment of inertia and modulus, and we show that this geometry can be mapped to a simpler system of two torsional springs as shown in Fig. 5(b). In elastic beam theory, the bending angle from the base of a cantilevered beam to its tip (Dh) scales linearly with the rotation angle at the free end (a) and can be expressed as af Dhf \u00bc bf ; am Dhm \u00bc bm. Geometrically, the [This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to ] IP: 149.150.51.237 On: Thu, 25 Sep 2014 22:49:39 ratios b must satisfy bm > bf > 1. (To provide some physical context for interpreting these geometric parameters, we note that bf and bm for a simple cantilever beam are 1",
            ") The compatibility condition leads to the expression for the dimensionless moment at the contacting tip, mc \u00bc kmbf kf bm fn cos h\u00f0 \u00de ft sin h\u00f0 \u00de ; (6) where kf and km are the dimensionless angular stiffness of the seta as a whole in response to a moment mg at the seta root imparted by either a bending force or a moment, respectively, applied at the seta tip. Note also that geometrically kf must be larger than km. From the balance of moments, this gives the ratio of contact to root moments mc mg \u00bc kmbf kf bm kmbf : (7) For a seta modeled as a curved beam, the values of kf, km, bf, and bm can be found from elementary mechanics using the unit load method. However, with the imposed boundary condition of zero net contact rotation, then we can find an equivalent two-torsional-spring system as depicted in Fig. 5(b) with kc \u00bc kmbf bm bf and kg \u00bc kf bm kmbf bm bf that undergoes the same displacement and possesses the same moment ratio as the geometry-based model for a given loading condition. In this simpler system, Dh \u00bc mc kc \u00bc mg kg and mc mg \u00bc kc kg \u00bc g: (8) This means that the system response is described by an effective root stiffness kg and stiffness ratio g. Thus, rather than estimating kf, km, bf, and bm for seta shaped beams (which would contain much uncertainty), we determine the stiffness scale and ratio g that is optimal for the gecko"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001988_3dp.2014.0017-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001988_3dp.2014.0017-Figure4-1.png",
        "caption": "Figure 4. CAD model of a crescent wrench ( left ) and the resulting printed part ( right ).",
        "texts": [
            " 17 While the ceramic salts performed well for a larger variety of materials, sucrose proved to be the best candidate for the bronze alloy used in these experiments. The bronze base powder used was a fully\u00a0 alloyed bronze with chemical composition shown in Table 1 . The powder particle size has a distribution of 98.7 % \u2212 325 mesh, 1.3 % \u2212 200 / + 325 mesh as reported by the manufacturer. This research has demonstrated the ability of the SIS - metal process to be adapted to a low - cost, inkjet printhead - based machine. Bronze parts were fabricated successfully using the SIS - metal machine. Figure 4 shows the CAD model of a crescent wrench and the resulting functional part printed by the SIS - metal machine. The crescent wrench was fabricated in separate pieces and then manually assembled. The four pieces consisted of the body with fixed jaw, worm screw, pin, and adjustable jaw. Figure 5 shows the CAD model of an edited M \u00f6 bius strip, downloaded from the GRABCAD website, 18 and the corresponding part fabricated by the SIS process. The M \u00f6 bius is a single - piece design inspired by jewelry applications"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000049_robio.2013.6739433-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000049_robio.2013.6739433-Figure3-1.png",
        "caption": "Fig. 3. Illustration of angular offset for a foil.",
        "texts": [
            " The depth-sounding sensor is capable of operating under 4000 m water, with the accuracy of centimeter level. What\u2019s more, a CCD camera made by Sony is also been equipped. A. Motion analysis of actuated-flippers The turtle-like underwater vehicle obtain the swim capacity through making the four flippers moving back and forth. Since the sinusoidal signal can generate smooth vibration to the flippers, and make it convenient for the adjustment of the control parameters, we introduce it to control the vehicle. As shown in Fig. 3, the joints of the underwater vehicle do simple harmonic motion as below:   sin( )i i i i it A t      (1) In the equation, i ( (1, 2, 3, 4))i presents the order number of the flipper. ( ) i t is the target angle at the time t of flipper i, i  is the angular offset, i A is the vibration amplitude, i  is the vibration frequency, i  presents the phase position. The magnitude of the force generated by the flipper can be adjusted by changing the vibration amplitude and vibration frequency, and the direction of the force is determined by the angular offset of the flipper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002146_j.mechmachtheory.2014.04.012-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002146_j.mechmachtheory.2014.04.012-Figure5-1.png",
        "caption": "Fig. 5. Coordinate systems for the shaving machine with cutter assembly errors.",
        "texts": [
            " The traveling distance dp of the pin along the guideway is converted from the traverse distance zt of the pivot and it can be expressed as follows: dp \u00bc dh \u00fe zt\u00f0 \u00de cos\u03b8\u00fe dv sin\u03b8\u2212 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi dh \u00fe zt\u00f0 \u00de cos\u03b8\u00fe dv sin\u03b8\u00bd 2\u2212zt 2dh \u00fe zt\u00f0 \u00de q ; \u00f015\u00de parameters dv and dh are the vertical and horizontal distances between the pin and the pivot, respectively, and \u03b8 is the where guideway inclination angle with respect to the horizontal slide (Fig. 4). The auxiliary crowning mechanism's crowning angle \u03c8t is dependent on the guideway inclination angle \u03b8 and the traverse distance zt. It can therefore be determined by \u03c8t \u00bc arctan dv dh \u2212 arctan dv\u2212dp sin\u03b8 dh\u2212dp cos\u03b8\u00fe zt ! : \u00f016\u00de If the rotation center of the guideway is adjusted to the same horizontal level of pivot, then when the angle of inclination \u03b8 of the guideway equals zero, the crowning angle \u03c8t also equals zero. Fig. 5 shows the schematic coordinate systems for the parallel gear shaving process with an auxiliary crowning motion and shaving cutter assembly errors. Herein, the coordinate systems S1(x1,y1,z1), S2(x2,y2,z2), and Sd(xd,yd,zd) are rigidly connected to the shaving cutter, work gear, and shaving machine's frame, respectively. Coordinate systems Sa(xa,ya,za), Sb(xb,yb,zb), Sc(xc,yc,zc), Se(xe,ye,ze), Sf(xf,yf,zf), and Sg(xg,yg,zg) are the auxiliary coordinate systems for the homogeneous coordinate transformation, while Sh(xh,yh,zh) and Sv(xv,yv,zv) are the auxiliary coordinate systems for determining the effects of shaving cutter assembly errors in the horizontal and vertical directions",
            "4 \u03bcm, which may cause vibration and noise when the work gear pair changes its rotational direction. Shaving cutter assembly error in the vertical axial direction \u0394vs induces a dislocation of the center cutting point between the shaving cutter andwork gear along the zd-axis. This dislocation is determined at the shortest center distance Eos between the shaving cutter andwork gear when zb-axis is parallel with zv-axis thatmeans that the vertical axial assembly error \u0394vs equals crowning angle (\u03c8t(zt)), as shown in Fig. 5. Thus, parameter a expressed in Eqs. (24) and (25) can be applied. After solving the equation \u0394vs = \u03c8t(zt) under the vertical axial assembly error of \u0394vs = 0.05\u00b0, the center cutting point dislocation zt equals 3.54 mm. This dislocation most affects tooth flank twist in the shavedwork gear, and themaximum flank twist increases from 13.8 \u03bcm(Fig. 7) to an extreme amount rameters for the double-crowned helical pinion 2, standard helical gear 3, and shaving cutter. of 60.8 \u03bcm (Fig. 9). If the shaving cutter has a horizontal axial assembly error (\u0394hs = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002867_arso.2016.7736297-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002867_arso.2016.7736297-Figure1-1.png",
        "caption": "Figure 1. Flexible robotic polishing system",
        "texts": [
            " Belt sander had a superior grinding performance, could adapt to the complex forming surface, and had certain flexibility. But the contact force during operation could not be controlled and the consistency of the material removal could not be ensured. For the workpiece difficult to hold, using grinding wheel and force feedback device mounted on the end of industrial robot for effective path planning could have the advantage of belt sander and ensure the consistency of the material removal by controlling polishing force. Therefore, the proposed flexible robotic polishing system was shown in Figure 1(a), including six-axis industrial robot A, the force feedback device B, electric sander C, workpiece to be machined - blade D, Manual turntable E and other components. Industrial robots clamped the electric sander by connecting force feedback device, processing the polishing operations on the workpiece fixed in the manual turntable. FIG. 1(b) was the robotic polishing apparatus proposed before, wherein the force feedback device B was fixed to the end of the robot A through adapting piece F, polishing wheel G was fixed in electric sander C and the force feedback device B connected to the electric sander C. The control system consisted of the 978-1-5090-4079-7/16/$31.00 \u00a92016 IEEE 289 industrial robot control system and the force feedback device control system. Robot path planning is one of the key technologies of flexible robotic polishing system",
            " Usually, the size of the belt could be regarded as the polishing process width because of the flexibility of the belt in the abrasive belt polishing process. In this paper, the polishing wheel was used to carry out the work. In the case of surfaces polishing, there was a need for analysis of the contacting deformation. The fit width between the wheel and the workpiece could be obtained afterwards, and it could provide the data to support for the polishing path planning. The contact model in this paper was mainly composed of two parts: the workpiece to be processed (as shown in Fig.1) and the grinding and polishing wheel. Blade was one of the most important components of the aero-engine. The experiment was based on the middle part of easy processed area by measuring the radius of curvature. The blade was equivalent to a sphere model for contact analysis through data processing. The horizontal line cutting method was adopted taking into account that the transverse curvature was the main factor that affects the width of the contact surface. We used the three-dimensional software to make a test of the horizontal curve curvature, and the results were as follows: The physical model was built by using ANSYS, and the polishing wheel was meshed into elements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003816_s12541-019-00014-2-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003816_s12541-019-00014-2-Figure2-1.png",
        "caption": "Fig. 2 Pressure distribution cloud map of the propeller blade",
        "texts": [
            " These masses of the parts and position of the center of mass are constant values if the structure of the assembly does not change. Then, only the rotation speed can change in the actual working conditions. In this paper, the rotation speed is set as RP , and the centrifugal forces with different rotation speeds are listed in Table\u00a01. In this work, the initial design model of the 3500KW CPP is regarded as a research object, and the diameter of propeller blade is 3\u00a0m and the number of the blades is 4 with right-handed rotation. Figure\u00a02 presents the pressure distribution cloud map of the propeller blade, and the component of the hydrodynamic force is listed in Table\u00a02 at different rotation speeds and inflow velocities based on the hydrodynamic simulation experiments. In Table\u00a02, the inflow velocity refers to the flow rate of the seawater around the CPP undersea. As shown in Fig.\u00a03, component (1) \u2211 Fz = 0 \u21d2 Frad = \u2212Fce,z \u2212 Fhd,z force Fhd,z is very small, accounting for 4% of the centrifugal component force Fce,z . In conclusion, the changes in the external conditions such as the inflow velocity affect the load of the blade bearing slightly, basically negligibly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003379_0954410018764472-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003379_0954410018764472-Figure3-1.png",
        "caption": "Figure 3. Coordinate axis system.",
        "texts": [
            " Thereby, the dynamic of the morphing motion is described by using Kane method in this paper. We suppose that the morphing aircraft consists of five separate rigid bodies: fuselage(Body1), inner part of left wing(Body2), outer part of left wing(Body3), inner part of right wing(Body4), and outer part of right wing(Body5), and their masses are mb,m1,m2, m1,m2, respectively. The important parameters of the morphing aircraft are given in Appendix 1. Firstly, the appropriate coordinate systems must be created in order to set up the Kane equations. As shown in Figure 3, we set the origin of the whole aircraft body coordinate frame Obxbybzb to locate at the center the wing (the unit vectors corresponding to the axes xb, yb, and zb are ebx, e b y and ebz), and the ground coordinate system is described as Ogxgygzg. The coordinate frames of morphing aircraft\u2019s fuselage and the other several moving parts of wing are described as Oixiyizi\u00f0i \u00bc 1, . . . 5\u00de, taking the center of mass of each body as the origin correspondingly. CG denotes the center of mass of the whole aircraft and dCG represents the distance between CG and the nose of the aircraft",
            " b the distance from Ob to O1, m CD ,CL, the coefficients of drag, lift, and CM moment CL \u00bc0 ,CL , lift force coefficient with respect to CL e \u00bc 0, , e CM \u00bc0 , pitching moment coefficient with CM ,CM e respect to \u00bc 0, , e D,L,T drag, lift, thrust, N g acceleration due to gravity m/s2 h altitude, m l1, l2 fixed length of inner wing and outer wing, m l the distance from Ob to O3 or O5, m Jx, Jy, Jz moments of inertia of the fuselage, kg/s2 J2, J3, moments of inertia of the Body J4, J5 i, kg/s2 mt the whole mass of aircraft, kg M pitching moment, N m q pitch angular velocity, rad/s Sw wing area, m2 T T throttle coefficient u,w scalar components of speed in body axes, m/s V velocity, m/s angle of attack, rad e elevator angular deflection, rad T throttle, % expansion of wing span, m l, scheduling parameters air density, kg/m3 pitch angle, rad sweep angle, deg Appendix 1. Important parameters of the morphing aircraft b 0:3 m l2 2:0 m 0 45 deg mb 771:79 kg 0 2 m m1 54 kg Jy 3107:53 kg m2 m2 13:605 kg l1 2:5 m T T 129:27 N=% Appendix 2 The specific definitions of variables in steps (2) and (3) of Figure 3 The generalized active force Fi Ob of body i can be given by F1 Ob \u00bc \u00f0T mbg sin \u00fe L sin D cos \u00deebx \u00fe \u00f0mbg cos D sin L cos \u00deebz F2 Ob \u00bc m1g sin e b x \u00fem1g cos e b z F3 Ob \u00bc m2g sin e b x \u00fem2g cos e b z F4 Ob \u00bc m1g sin e b x \u00fem1g cos e b z F5 Ob \u00bc m2g sin e b x \u00fem2g cos e b z 8>>>>>>>< >>>>>>>: The generalized active moment Mi Ob can be given by M1 Ob \u00bcMeby M2 Ob \u00bcM3 Ob \u00bcM4 Ob \u00bcM5 Ob \u00bc 0 ( The partial velocities U i\u00f0 \u00de Ob and W i\u00f0 \u00de Ob of each body i are defined as U i\u00f0u\u00de Ob \u00bc ebx, i \u00bc 1, 2 . . . 5 U i\u00f0w\u00de Ob \u00bc ebz , i \u00bc 1, 2 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure5-1.png",
        "caption": "Fig. 5. Cylindrical body in the water\u2013vapor flow.",
        "texts": [
            " To model the heat transfer temperature of an outside wall is considered to have constant temperature and equal to 150 C. Figs. 3 and 4 show the drag force changes at the surface of the cold solid cylindrical body. It can be noticed that the highest velocity is at the front part of the edges of the body where water flow is forced to rebound. The highest force is at the front and at the end part of the body. Total drag force, which influences the motion of the body, is 0,015 N. Due to the vortex beyond the cylindrical body, the drag force at the end part becomes negative. Fig. 5 shows water velocity change nearby the hot body. One can see in Figs. 3 and 5 s a significant difference on the front zone where the flow abruption appears. Fig. 6 shows the change of flow density. Sharp decrease in density appears nearby the cylindrical surface. Fig. 7 shows a change of vapor mass fraction on the surface of the body. Maximum vapor generation is noticed on the front part of the cylinder immediately behind the flow abruption point. Average vapor mass fraction on the surface is 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002009_j.precisioneng.2014.02.012-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002009_j.precisioneng.2014.02.012-Figure4-1.png",
        "caption": "Fig. 4. Linear motion rolling bearing.",
        "texts": [
            " As the ratio increased, he resultant velocity vector came close to the vibration direction or an increasing proportion of the vibration cycle, resulting in a eduction in the measured frictional force. . Elements for ultrasonic oscillation The current study was conducted by driving a single-axis cariage in a sinusoidal motion as a simulation of circular motion, as roposed by Tanaka et al. [36]. Ultrasonic oscillation was applied to he rail and carriage, and their effect on the frictional force during otion was investigated. Fig. 4 shows the commercially available linear motion rolling bearing used in this experiment. The rolling bearing consisted of a carriage (IKO Nippon Thompson Co., Ltd., LWH30) and a rail (IKO Nippon Thompson Co., Ltd., LWH30 R160), which used bearing balls as the rolling element. When the carriage moved along the rail, the bearing balls rubbed the raceways and generated rolling friction. To reduce the friction force, each raceway had to be excited by an ultrasonic actuator, since it was impossible to excite the bearing balls alone"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003609_s1061934818100039-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003609_s1061934818100039-Figure4-1.png",
        "caption": "Fig. 4. A schematic diagram of the biosensor.",
        "texts": [
            " The potential of cyanidase (cyanide dihydratase) ANALYTICAL CHEMISTRY Vol. 73 No. 10 2018 lized whole cell cyanide dihydratase of F. indicum (2). 0.5 0.7 0.9 1.1 2 1 /mM for the construction of biosensors has been suggested by Keusgen et al. (2001) [25] based on flow through apparatus with a linear range 0.6\u201330 ppm (23 \u03bcM\u2013 1.1 mM) and a detection limit of 0.2 ppm to measure formate. In the present study, a potentiometric approach was followed to design a biosensor for instant detection of cyanide (Fig. 4). Whole cell cyanide dihydratase of F. indicum has been used as a biocomponent for this purpose. Experiments were carried out to optimize enzyme loading for maximum potential response. It can be observed from Fig. 5 that an increase in potential response was recorded with the increase in the amount of enzyme loading upto 0.3 U/mg dcw enzyme units and beyond that a decrease in potential response was measured. The optimized enzyme quantity (0.3 U/mg dcw) was used in further experiments. The response time for different cyanide concentrations was also measured"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003620_j.matpr.2018.06.241-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003620_j.matpr.2018.06.241-Figure6-1.png",
        "caption": "Fig. 6 Stress results for loading cases (iii) and (iv)",
        "texts": [
            "/ Materials Today: Proceedings 5 (2018) 18920\u201318926 18925 The graphs show the average stresses induced in different tubes in MPa. It can be seen that in all the cases the seat stays are comparatively stresses more followed by the top tube. The seat tube is the least stressed in all the cases. It can be seen that the maximum stresses are located at the junction of the tubes and stress concentration occurs. Also, all the stresses are well below the yield stress of the material used as shown in above Fig 6 (i-iv). (i) (ii) 18926 Devaiah B.B et al./ Materials Today: Proceedings 5 (2018) 18920\u201318926 In the case where the rigid links were used to simulate real world scenario\u2019s the fork seems to be getting stressed more and frame seems to be unaffected. Therefore, a static structural analysis wasn\u2019t conclusive. A further exhaustive study must be done to debug this problem as shown in above Fig 7 (i-iii). In the present the finite element analysis of a Bicycle frame was carried out using ANSYS FEM software using various boundary conditions and the results were compared with theoretical results found in the literature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001215_iscas.2017.8050981-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001215_iscas.2017.8050981-Figure1-1.png",
        "caption": "Fig. 1. Android-Based Robotics Platform. The robot runs on a Dagu Wild Thumper 6-wheel-drive all-terrain chassis, with an SPT 200 pan and tilt to hold the Samsung Galaxy S5 smartphone and control the view of the phone camera. Front-facing MaxBotix LV-MaxSonars can detect obstacles. An ION Motion motor controller and IOIO-OTG microcontroller are housed in the back of the robot. Computing is handled by the Android phone, which accesses the sensors and actuators through a Bluetooth connection with the IOIO-OTG.",
        "texts": [
            " In this paper, we describe our current advances in developing an integrated neuromorphic system for performing all aspects of outdoor navigation. 1) Android-Based Robotics Platform: In creating an integrated neuromorphic system, we needed to create an experimental setup in which long term and reactive motion planning could be isolated and swapped for neuromorphic or traditional implementations. The Android-Based Robotics Platform was created at the University of California, Irvine, consisting of commercial off-the-shelf components [7]. A durable ground robot controlled by an Android smartphone (Figure 1) was ideal for an outdoor neuromorphic system of navigation, as it can run simulations of neuromorphic algorithms and has all of the necessary hardware for communication and localization. Instructions for building the robot can be found at http://www.socsci.uci.edu/\u223cjkrichma/ABR/. With this platform we were able to test the robot in Aldrich Park, a 19-acre botanical garden at the University of California, Irvine, which contained roads of different widths, terrains, and inclinations. 2) Spiking Wavefront Propagation and Robot Implementation: We created an algorithm for path planning that could be run on neuromorphic hardware [8]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000097_educon.2012.6201108-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000097_educon.2012.6201108-Figure3-1.png",
        "caption": "Fig. 3. Servo motors, X, Y, Z tilt sensor, control boards and video camera top view",
        "texts": [
            " Internet Workshop Server / Database FileMaker Instant Web Publishing commands / requests readings / data video Bi-directional converter TCP / IP - logic LAN Driver\u2019s Load AC/DC motor, Stepper, Servo, rectifier, converter, etc. Sensors I, U, MEMS, Hall effect, etc. Video camera USB or LAN Fig.1. e-Learning workshop structure. III. DRIVERS AND LOADS The main goal of the e-Workshop is to give possibility for students to have semi-vocational practice on basic parameter relationships, control methods and algorithms, as mentioned above. Two servo drive system with installed MEMS tilt sensor (based on MEMS accelerometer) is shown on Figure 2. Servo drive system from top view is shown on Figure 3. User control interface of e-Workshop (web browser screen-shoot) is shown on Figure 4. Systems two RC servo motors are digitally controlled from Arduino - Arduino Ethernet boards. Servo motor max are +/- 90 deg against middle position, set by sending 1,5ms long control signal every 20ms. Servo motor axle angular position changes are managed by changing impulse length. Servo motors axis are positioned on right angle and 2nd servo is mounted on 1st servo axle. MEMS sensor board is mounted on 2nd servo motor axle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003335_09544828.2018.1435853-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003335_09544828.2018.1435853-Figure8-1.png",
        "caption": "Figure 8. Case study\u2019s main body components and roof before assembly, courtesy of IKKCO.",
        "texts": [
            " It is to be noted that for each sampling point, an FEA approach followed by an MVS fairness computation is needed. However, it is a clearly reasonable approach comparing toMonte Carlo simulation, which demands a larger population to achieve desirable results. The overall process layout of the proposedmethodology is demonstrated in Figure 7. In this section, the surface fairness metric of the car\u2019s roof, as one of the key quality characteristics in automotive bodies, is statistically evaluated in the tolerance analysis procedure. Figure 8 illustrates a car in which the assembly process of the roof is the last stage of the body assembly line and, therefore, all accumulated deviations from the previous stations should be defined as input tolerances. Figure 9 only shows those parts that have the most effect on the surface variation of the roof. Other components are excluded from themodel and replaced with appropriate constraints in order to increase the analysis speed. For simplification, it is presumed that the deviation of left and right side bodies from their nominal positions (Figure 10) is the only input variation of the analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001294_tpel.2017.2782804-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001294_tpel.2017.2782804-Figure2-1.png",
        "caption": "Fig. 2. Flux linking between a \u2212 a\u2032 and x \u2212 x\u2032 coils: (a) maximum when \u03b8 = \u03c0 2 + \u03c0 12 , and (b) minimum when \u03b8 = \u03c0 12 .",
        "texts": [
            " Then the flux linkage of abc coil is described by \u03bbabc = Liabc + Mixyz \u2212 L\u03b4A(\u03b8)iabc \u2212 L\u03b4A(\u03b8 \u2212 \u03c0 12 )ixyz +\u03c8(\u03b8), (11) where M = \u221a 3Lm 2  1 \u22121 0 0 1 \u22121 \u22121 0 1  . (12) Note that M is a static contribution to \u03bbabc made by the xyz-coil shifted by \u03c0/6, and that the matrices, [Labc : M] already appeared in six phase SPMSMs [27], [28]. The PM contribution \u03c8(\u03b8) is also the same as that of SPMSM. The reluctance part of the self inductance, \u2212L\u03b4A(\u03b8) is the same as in the case of three phase IPMSM. However, we have the other reluctance component from the xyz-coil, and it is described by \u2212L\u03b4A(\u03b8\u2212 \u03c0 12 ). Fig. 2 illustrates two extremes: The flux linking from x\u2212x\u2032 to a\u2212 a\u2032 is maximized when the PM center is aligned to the middle point of a and x coils, since the PM location gives a minimum hindrance to the flux as shown in Fig. 2 (a). The rotor angle is \u03b8 = \u03c0 2 + \u03c0 12 , thereby A(\u03b8 \u2212 \u03c0 12 ) \u2223\u2223 \u03b8=\u03c0 2 + \u03c0 12 = A(\u03c02 ). Note that \u2212L\u03b4A(\u03c02 )(1,1) = \u2212L\u03b4 cos\u03c0 = L\u03b4 yields the maximum value. The other case is shown in Fig. 2 (b) in which the PM location gives a maximum hindrance to the flux. Specifically, the flux linking from x \u2212 x\u2032 to a \u2212 a\u2032 is minimized when \u03b8 = \u03c0 12 . Then, A(\u03b8 \u2212 \u03c0 12 ) \u2223\u2223 \u03b8= \u03c0 12 = A(0). Note that \u2212L\u03b4A(0)(1,1) = \u2212L\u03b4 cos 0 = \u2212L\u03b4 yields the minimum value. Therefore, the mutual inductance matrix between abc and xyz coils are modeled by A(\u03b8 \u2212 \u03c0 12 ) for arbitrary angle, \u03b8. Let the flux linkage of xyz-coil denoted by \u03bbxyz = [ \u03bbx, \u03bby, \u03bbz ]T . Following the similar reasoning, it follows that \u03bbxyz = Lixyz + MT iabc \u2212 L\u03b4A(\u03b8 \u2212 \u03c0 12 )iabc \u2212 L\u03b4A(\u03b8 \u2212 \u03c0 6 )ixyz +\u03c8(\u03b8 \u2212 \u03c0 6 )"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003129_issc.2017.7983613-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003129_issc.2017.7983613-Figure1-1.png",
        "caption": "Fig. 1. Ship motion - six degrees of freedom",
        "texts": [
            " For simulation purposes and to prove the concept only a five second and a one second landing training was undertaken. As the UAV approaches the deck the landing time will decrease. To deal with this ANNs could be trained for four, three and two second landings. The landing time is a function of the distance between UAV and the landing pad. As shown earlier the trained ANN from Phase 2 can calculate this distance, so calculations can be made to estimate the landing time from a specific height above the landing pad. III. RESULTS Matlab and Unity were compared for corner deducing ability. Fig. 1 below depicts a visual comparison between two orientations, the left side shows Matlab\u2019s Harris corner recognition algorithm results and the right side shows Unity\u2019s camera window size used to calculate the screen co-ordinates of the \u201cH\u201d. Ten images of random ship orientations were fed into Matlab and the corner recognition results were compared with Unity\u2019s calculations. The average difference between both sets of results was 0.3%. The ANN described earlier for Phase 2 trained quickly and yielded a high regression of 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002322_j.triboint.2015.10.007-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002322_j.triboint.2015.10.007-Figure2-1.png",
        "caption": "Fig. 2. Photograph of the water-sealing system.",
        "texts": [],
        "surrounding_texts": [
            "2.1. Configuration of the water-sealing system Figs. 1 and 2 show the configuration of the proposed watersealing system, implemented as a shaft seal. Hydrated seal lips made of PVF were attached to a rotating shaft. The rotating shaft had a diameter of 30 mm and was fabricated from stainless steel (SUS304, JIS). The dynamic seal face on the rotating shaft was polished with a surface finish of approximately 0.02 \u03bcm (Ra). Fig. 3 shows a three-dimensional optical surface profile of the shaft (NewView 7000, Zygo Corp., USA). Previous studies have revealed that the surface finish of the shaft of an oil seal should be strictly controlled; furthermore, other parameters, such as skewness (Rsk), kurtosis (Rku) and the maximum depth of the surface profile below the mean line (Rv), have been suggested to influence oilseal performance [13]. In this study, a simple polishing method using alumina particles of 1.2 \u03bcm (WA #8000) was adopted, and no specific control of the surface finish was applied. An aqueous solution of PEG, which is a non-Newtonian fluid, was injected between the two seal lips for lubrication and sealing action at the dynamic seal faces. This configuration allowed the shaft to move in the direction of the long axis. A thrust bearing was implemented to prevent movement and journal bearings to support the shaft."
        ]
    },
    {
        "image_filename": "designv11_13_0001038_asjc.1502-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001038_asjc.1502-Figure1-1.png",
        "caption": "Fig. 1. TRMS laboratory unit [1]. [Color figure can be viewed at wileyonlinelibrary.com]",
        "texts": [
            " As a consequence, TRMS presents higher coupling between dynamics of the rigid body and that of the rotors as compared to a conventional helicopter, and yields highly non-linear, strongly crosscoupled dynamics. TRMS has one main rotor and one tail rotor with fewer degrees of freedom of movement and more involved plant dynamics as compared to a conventional rotorcraft. Thus, this MIMO plant may be tested in a laboratory for its controllability in two degrees of freedom of movement through two degrees of freedom of control. Hence, this set-up is also a laboratory standard for experiments and to understand the effect of cross coupling in 2 \u00d7 2 systems in specific and MIMO systems in general. Fig. 1 illustrates a simple schematic diagram of TRMS at rest. Determination of the nominal model of a physical plant is the first step in any control system problem in general. If the required nominal model is linear, this could be determined either by linearizing the mathematical relationship of the system and obtaining the state space model or by system identification/dynamic modeling. Several efforts have been made in the past to derive a representative linear model of TRMS that behaves in the same manner as the physical TRMS in hover (translatory motion w"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001585_aim.2018.8452392-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001585_aim.2018.8452392-Figure10-1.png",
        "caption": "Figure 10. Unit part of proposed robot",
        "texts": [
            " The robot was also problematic in that the amount of contraction in the unit part alone was small, and the air pressure response was poor. The robot was difficult to handle when put to practical use. A peristaltic motion-type robot was developed with the aim of improving the motion of the robot with a view toward practical applications and improving basic performance. The authors developed another peristaltic motion robot named PEW-RO V. A schematic of PEW-RO V is shown in Fig.8. The specifications of the robot are shown in Table 3. Fig.10 shows the contraction of the unit part in the pipe. The amount of contraction inside the pipe depends only on the diameter of the artificial muscle and the diameter inside the pipe (Eq. (1)). Compared with the previous unit part, the diameter of the artificial muscle is thinned by reducing the size of the artificial muscle fastener (Fig.11). As a result, it was confirmed that the contraction inside the tube improved by 70%. The contraction characteristics of the existing robot unit are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003356_j.apm.2018.01.018-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003356_j.apm.2018.01.018-Figure8-1.png",
        "caption": "Fig. 8. Boundary conditions for the loading analysis.",
        "texts": [
            " Again, a deformable-to-rigid (Meshed-to-Geometric) touching contact table was specified for the meshed tyre and the geometric road surface contact interface. A friction coefficient of 0.7 was used to model the friction between the tyre and road. This coefficient is typically used for gravel road contact [19] . After the model was inflated, a Point Load was applied to a control node located on the road surface to transfer the load to the rigid road surface. The tyre rim was kept fixed and the load was applied. The boundary conditions used for the 3D model and road surface are shown in Fig. 8 . The vertical tyre displacement obtained from the loading analysis was used to load the tyre in the rolling analysis using a displacement control condition. The 3D tyre model for the loading analysis were used for the rolling analysis as well. However, for the rolling analysis, the tyre was rotated 180 \u00b0 around the axial axis. The road surface was extended to a length of 14 m. The rolling model and road surface are shown in Fig. 9 . The numerical vertical tyre displacement was applied to the road surface using a Velocity Control Rigid Body condition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003423_physrevfluids.3.043101-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003423_physrevfluids.3.043101-Figure6-1.png",
        "caption": "FIG. 6. Similar to Fig. 5 except that Pe = 500 and (a) U1/U0 = \u22120.042, (b) U1/U0 = 0.033, and (c) U1/U0 = \u22120.012.",
        "texts": [
            " 5(b), because of the symmetric motion of the swimmer surface and nearly symmetric distribution of c with respect to the y axis except near \u03b3 = \u00b11, the contour of cm around the swimmer is symmetric. In this case, negative regions grow in size and result in a larger increase of U1/U0 than in the similar case at low Pe. From Fig. 5(c) it can be seen that in the case of \u03b24/\u03b21 = 1 the dominant negative region of cm switches from the back of the swimmer to its front at \u03b3 \u223c 0.6 and the region with positive cm increases in both size and magnitude, resulting in small U1/U0. At very high Pe, Pe = 500 in Fig. 6, the main contribution of the c distribution only concentrates at the boundary layer of the swimmer and the largest value of cm, either positive or negative, happens along the surface of the swimmer. The exception is at isolated stagnation points along the surface motion where the nutrient boundary layer separates. The stagnation points are placed near the back of the squirmer where a larger region with small magnitude of negative cm bursts from the surface and stretches into the wake. The effective consequence of an increase in the strength of the region with positive cm and detachment of the region with negative cm from the surface of the swimmer (where \u03c3 \u2032 has a larger value) is a reduction in U1/U0 compared to the smaller Pe regime of Pe = 50"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000969_j.jsv.2017.01.018-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000969_j.jsv.2017.01.018-Figure7-1.png",
        "caption": "Fig. 7. Experimental apparatus.",
        "texts": [
            "018i For the analysis of the wiper blade during reversal, we have developed a set of simulation algorithms using augmented methods and Baumgarte's stabilization method and the derivation method to extract the exact transition times using slack variable method. Using these simulation algorithms, the reversal behavior of the wiper blade is clarified theoretically. In this section, we present the experimental results of the reversal behavior of an actual wiper blade. For the experimental apparatus (Fig. 7), a cut-off segment of a wiper blade was mounted on a linear actuator that swept it across a plate glass surface. The holder with the wiper blade mounted at its tip is attached to the linear bearing and can move freely in the vertical direction. A spring of constant rigidity k0 with a screw tensioner mounted on the holder apply a compressive force to the holder. The initial compressive force can be changed by adjusting the screw and providing the initial displacement hd to the spring. These mechanisms are fixed on the linear actuator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003307_6.2018-0449-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003307_6.2018-0449-Figure3-1.png",
        "caption": "Figure 3. Example of deployed membrane employing rotational spring.",
        "texts": [
            " A nonlinear rotational spring is introduced to the crease of the membrane in order to express the plastic deformation of the crease and to reduce the number of nodes around the crease. The properties of the rotational spring are obtained by the results of the deployment simulation in Sec.V, which is represented by the deployment angle  and the tensile force PD. The deployment angle  in the folded configuration is zero in order to determine initial shape of the membrane, and thus, the tensile force at  = 0 deg becomes negative value as shown in the schematic drawing of the  \u2013 PD relationships in Fig.2. Figure 3 shows an example of deployed membrane by employing the rotational spring. The rotational spring in Fig.1 and 2 is treated by connector element in finite element program, ABAQUS. The employment of the connector element has two positive aspects. The first one is reduction of the number of elements. The other is improvement of the accuracy of the crease behavior because the shell element is difficult to treat large strain in the thickness direction even if large number of elements are used. In order to express the nonlinear crease behavior, the following three options are employed to express the crease"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure5.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure5.1-1.png",
        "caption": "Fig. 5.1 Cylindrical model of the induction machine magnetic circuit (a) simplified scheme of the induction machine magnetic circuit; (b) induction machine magnetic circuit as a multi-layer structure",
        "texts": [
            " 4, it was shown that the regions of the magnetic circuit can be represented as layers with smooth surfaces. The regions reflecting the corresponding stator and rotor teeth areas can be replaced by layers with average values of magnetic permeability. In this connection, the magnetic circuit model can be considered as a multi-layer structure. If one proceeds from the fact that the magnetic circuit includes the air gap, stator, and rotor slot wedge, the wound parts of the teeth, and the joke regions (Fig. 5.1\u0430), then the magnetic circuit model can be obtained as shown in Fig. 5.1b, where the selected magnetic circuit regions are represented as coaxial cylindrical layers. The stator and rotor teeth regions include the axial slots (Fig. 5.1\u0430). The presence of the slots on the stator and rotor surfaces gives rise to the fact that the regions of the teeth areas have different reluctance values for the radial and tangential \u00a9 Springer International Publishing Switzerland 2015 V. Asanbayev, Alternating Current Multi-Circuit Electric Machines, DOI 10.1007/978-3-319-10109-5_5 127 components of magnetic flux",
            " To take into account this factor, the calculated length of the air gap is used in the technical literature. The calculated length of the air gap is determined using the well-known Carter\u2019s factor. The regions reflecting the stator and rotor joke areas are replaced by the layers with magneto-isotropic properties. For the stators of large electric machines, coldrolled steel is used. In this case, the stator joke region is presented as a magnetoanisotropic layer. On the basis of the model shown in Fig. 5.1b, a description of the field distribution in the magnetic circuit of an electric machine can be implemented by field equations, for example [1\u20133]. The solutions of these equations are presented in (4.69), and we use them to determine the magnetic field components in the selected magnetic circuit regions. In the model in Fig. 5.1b, the surface 0 0 passes through the middle of the air gap, and it divides the electric machine magnetic circuit into two systems of cylindrical layers. The system of outer cylindrical layers (in relation to the surface 0 0) represents the stator magnetic circuit, and the system of inner cylindrical layers reflects the rotor magnetic circuit. We assume that the normal component of magnetic induction is given on the surface 0 0 (Fig. 5.1b). First, we consider the system of outer layers representing the stator magnetic circuit. In Fig. 5.1b, the stator magnetic circuit is divided into four layers, which represent the stator joke, slot wedge, wound part of the tooth regions, and the upper half of the air gap. The equations for the magnetic field components in the stator magnetic circuit regions can be obtained by (4.69). We will assume that a relatively small portion of the magnetic flux is pushed aside from the upper surface of the stator joke region, and therefore can be neglected. In this connection, a consideration of the magnetic field in the stator magnetic circuit should start with a description of the magnetic field in the stator joke region",
            "34) acquire H0 \u00bc A0 1=R0\u00f0 \u00de Rp \u00fe R2p 01M01R p sin p\u03c6 \u00f05:38\u00de B0 \u00bc A0 \u03bc0=R0\u00f0 \u00de Rp R2p 01M01R p cos p\u03c6 \u00f05:39\u00de Equations (5.38) and (5.39) can be used to describe the values of the magnetic field components in the air gap. Below, we use the equations shown in (5.9), (5.10), (5.18), (5.19), (5.28), (5.29), (5.38), and (5.39) to construct the equivalent circuits for the air gap and stator magnetic circuit regions. We consider the system of equations (5.38) and (5.39) to construct the equivalent circuit for the air gap. In the model shown in Fig. 5.1b, the air gap is subdivided into two identical sub-layers. The upper half of the air gap reflects the stator magnetic circuit, and the lower half belongs to the rotor magnetic circuit. We will define the equivalent circuits of these air gap sub-layers. For this purpose, we will first 5.3 The Air Gap: The Equivalent Circuits 135 construct the equivalent circuit for the air gap layer with a thickness equal to \u03b4, for which we use the equations given in (5.38) and (5.39). Then, from the air gap equivalent circuit obtained in this way, we will define the equivalent circuits for its two sub-layers",
            "49) can be reduced to the form E01 \u00bc \u03c91\u03bcRksk\u03bcks\u03beLksR0 p \u03b8\u0446ksH01 \u00bc xkzaH01 \u00f05:50\u00de where xkza \u00bc \u03c91\u03bcRksk\u03bcks\u03beLksR0 p \u03b8\u0446ks Now, by equations (5.43) and (5.50), the equivalent circuit in Fig. 5.2 can be reduced to the form in Fig. 5.3. In accordance with Fig. 5.3, the air gap is replaced by the T-circuit. The reactance values of this equivalent circuit are determined by the expressions given in (5.45). In relation to the surface passing through the middle of the air gap (R\u00bcR0), the air gap layer is represented as subdivided into two identical sub-layers (Fig. 5.1). As was noted earlier, the upper sub-layer of the air gap reflects the stator magnetic circuit, and its lower sub-layer is included in the rotor magnetic circuit. We present these sub-layers in the form of corresponding equivalent circuits. For this purpose, 5.3 The Air Gap: The Equivalent Circuits 137 we use the T-circuit of the air gap in Fig. 5.3. The reactance of the transverse branch 0 0 of this T-circuit can be represented as the total reactance of two identical parallel reactance values equal to x01\u00bc x02\u00bc 2x0",
            "77) we have E 0 as1 \u00bc x\u03c4asc 2 ksc 2 zs H 00 as1 \u00f05:78\u00de For the values E 0 as1 and H 0 0 as1 used in (5.78), analogous with (5.76), it follows that E 0 as1 \u00bc Eas1cksczs andH 0 0 as1 \u00bc Has1 cksczs \u00f05:79\u00de Equation (5.78) represents the stator joke region. This equation describes a two-terminal network, and thus the stator joke region is replaced by the equivalent circuit of a two-terminal network (Fig. 5.11). In the equivalent circuit in Fig. 5.11, the value of x\u03c4as reflects the magnetizing reactance of the stator joke region. In accordance with the model in Fig. 5.1b, the stator magnetic circuit consists of four layers, which represent the stator joke, slot wedge, and wound part of the tooth regions, as well as the upper half of the air gap. The corresponding equivalent circuits for these regions have been obtained above. The L-circuit in Fig. 5.4 reflects the upper half of the air gap, the L-circuit in Fig. 5.8 represents the slot wedge region, and the L-circuit in Fig. 5.10 corresponds to the wound part of the tooth region. The equivalent circuit in Fig. 5",
            " On terminals ks ks of the equivalent circuits in Figs. 5.4 and 5.8, we have conditions E01\u00bcEks1 and as as Eas1 ' \"Has1 x\u03c4ascksczs 2 2 Fig. 5.11 One-port network representation of the stator joke region 5.5 The Modular Method: The Stator Equivalent Circuit 145 H01\u00bcHks1. Now, with consideration for these conditions, the equivalent circuits given in Figs. 5.4, 5.8, 5.10, and 5.11 can be connected in cascade. As a result, the equivalent circuit for the stator magnetic circuit obtains the form in Fig. 5.12. Thus, in accordance with the model in Fig. 5.1b, the equivalent circuit of the stator magnetic circuit follows as a result of the cascade connection of the L-circuits representing the upper half of the air gap, slot wedge, and wound part of the tooth regions, as well as the equivalent circuit for a two-terminal network reflecting the stator joke region. The equivalent circuit for the stator magnetic circuit obtained in this way (Fig. 5.12) represents a multi-loop circuit with a mixed connection of the elements (ladder network). The equivalent circuit in Fig",
            "94), radius Ras2 can be expressed as Ras2 \u00bc Ras1 \u00fe has \u00bc Ras1 1\u00fe 1 nas nashas Ras1 \u00bc Ras1mas, where mas \u00bc 1\u00fe 1 nas nashas Ras1 . With consider- ation for this condition, and taking into account that (nashas/Ras1)\u00bc nas(mas 1), for reactance x\u03c4as, from (5.94) we can receive x\u03c4as \u00bc \u03c91\u03bc\u03c6ashas 1 nashas=Ras1\u00f0 \u00de \u03beLas \u03beRas R2nas as2 R2nas as1 R2nas as2 \u00fe R2nas as1 \u00bc \u03c91\u03bc\u03c6ashas \u03beLas \u03beRas m2nas as 1 nas mas 1\u00f0 \u00de m2nas as \u00fe 1 \u00f05:95\u00de The analogous expression for the magnetizing reactance of the stator joke region was obtained by another method in Chap. 4. In accordance with Fig. 5.1b, the rotor magnetic circuit consists of four layers, which represent the rotor joke, slot wedge, wound part of the rotor tooth regions, and the lower half of the air gap. Thus the rotor magnetic circuit consists of regions analogous to the stator magnetic circuit. In this connection, it follows by analogy that the rotor equivalent circuit takes the form of the stator equivalent circuit shown in Fig. 5.12. The equivalent circuit for the stator magnetic circuit (Fig. 5.12) has been obtained as a result of the cascade connection of the L-circuits representing the upper half of the air gap, slot wedge, and wound part of the tooth regions, as well as a two-terminal network reflecting the stator joke region",
            "15 correspond to the outer surface of the rotor joke region, on which the conditions EzR2\u00bcEaR1 and HzR2\u00bcHaR1 are true. On the basis of expression (5.8), we receive for HaR1 HaR1 \u00bc RaR1 R0 H\u03c6aR1 \u00f05:115\u00de Here, EaR1 and HaR1 are the calculated values of the electric and magnetic field strengths on the surface of the rotor joke region. For the values of E 0 aR1 and H 0 0 aR1 used in the equivalent circuit in Fig. 5.15, by analogy with (5.79) we have E 0 aR1 \u00bc EaR1ckRczR andH 0 0 aR1 \u00bc HaR1 ckRczR \u00f05:116\u00de In accordance with the model in Fig. 5.1b, the rotor magnetic circuit consists of four layers, which represent the rotor joke, slot wedge, wound part of the tooth regions, and the lower half of the air gap. The corresponding equivalent circuits for these rotor regions have been obtained above. The selected regions of the rotor magnetic circuit are replaced by the corresponding L-circuits (with the exception of the rotor joke region, having the equivalent circuit of a two-terminal network). The L-circuit in Fig. 5.5 reflects the lower half of the air gap, the L-circuit in Fig",
            " On terminals zR zR of the equivalent circuits in Figs. 5.13 and 5.14, the conditions E 0 kR2 \u00bcE 0 zR1 and H 0 0 kR2 \u00bcH 0 0 zR1 are satisfied. On terminals kR kR of the equivalent circuits in Figs. 5.5 and 5.13, we have the conditions E02\u00bcEkR1 and H02\u00bcHkR1. Now, with consideration for these conditions, the equivalent circuits given in Figs. 5.5, 5.13, 5.14, and 5.15 can be connected in cascade. As a result, the rotor equivalent circuit takes the form in Fig. 5.16. Thus, in accordance with the model in Fig. 5.1b, the rotor equivalent circuit follows as a result of the cascade connection of the L-circuits representing the lower half of the air gap, slot wedge, and wound part of the tooth regions, as well as the equivalent circuit of a two-terminal network reflecting the rotor joke region. The rotor equivalent circuit obtained in this way (Fig. 5.16) represents a multi-loop circuit with a mixed connection of the elements (ladder network). The equivalent circuit in Fig. 5.16 is different from the equivalent circuit for the rotor magnetic circuit in Fig",
            " 5 on the basis of the cylindrical magnetic circuit models are represented in this chapter through corresponding planar model expressions. Expressions of magnetizing reactance values obtained in this way take into account the curvature of the surfaces of the magnetic circuit regions. As a result, simpler expressions for magnetizing reactance values of magnetic circuit regions are derived. This provision is realized in this chapter. A planar model of the electric machine magnetic circuit shown in Fig. 6.1b can be obtained from the cylindrical model in Fig. 5.1b by extracting its cylindrical layers into the planar layers in relation to the surface passing through the middle of the air gap and leaning on the radius R0. Derivation of the planar model in this way due to the fact that the expressions for the magnetizing reactance values of the magnetic circuit regions have been represented in Chap. 5 in relation to this air gap surface (R\u00bcR0). Cylindrical layers reflecting the stator regions are stretched into the planar layers in relation to their inner surfaces. These stator regions include the slot wedge, the wound part of the tooth and the stator joke areas. The cylindrical layers placed below the stator bore surface (including the air gap layer) are stretched into the planar layers in relation to their outer surfaces. In accordance with Fig. 5.1b, this area of the magnetic circuit contains the air gap, slot wedge, the wound part of the tooth and the rotor joke regions. \u00a9 Springer International Publishing Switzerland 2015 V. Asanbayev, Alternating Current Multi-Circuit Electric Machines, DOI 10.1007/978-3-319-10109-5_6 159 Extraction of the cylindrical layers (model in Fig. 5.1b) into the planar layers can be implemented in several steps. In the first step the ith cylindrical layer (Fig. 5.1b) is stretched into the ith planar layer with the pole pitch equal to \u03c4i. The value of the pole pitch of the ith planar layer \u03c4i is determined as \u03c4i\u00bc (\u03c0Ri1/p), where i\u00bc as, zs, ks, 0, kR, zR, aR. Here Ri1 is the surface radius (inner when Ri1 R01 and outer when Ri1 R01)of the ith cylindrical layer, where R01 is the radius of the stator bore surface (Fig. 5.1b). Consequently, a system of conditional planar layers having, in general, different axial lengths and different pole pitches can be obtained. Magnetic permeabilities (\u03bcxi and \u03bcyi) of the planar layers representing the corresponding stator and rotor teeth regions can be determined by the expressions given in (5.3). The corresponding expressions for the magnetic field components can then be written (in the Cartesian coordinate system). These expressions can then be reduced to the active length of the air gap l0 and to the pole pitch equal to \u03c40, where \u03c40 is the value of the pole pitch on the surface passing through the middle of the air gap",
            " When using numerical methods the magnetic field is represented as numerical data. In this case, on the surfaces of the air gap and stator and rotor regions, the numerical values of the components of the field strengths E and H (sinusoidally distributed 204 6 Magnetic Circuit Regions: Magnetizing Reactance Values in Terms of the Curvature within the pole pitch) can be established. In relation to these values of E and H, the electric machine magnetic circuit model can be represented as the multi-layer structure [2\u201310]. As such, we can proceed from the multi-layer model in Fig. 5.1b, on the surfaces of the corresponding layers of which the numerical values of the components of the field strengths E and H are given. For the sth layer, the equations composed in relation to the numerical values of E and H can be represented by analogy with (6.149) in the following form Es1\u00bc z11Hs1 \u00fe z12Hs2 Es2\u00bc z12Hs1 \u00fe z11Hs2 \u00f06:154\u00de Here the values of Es1, Hs1 and Es2, Hs2 are known and are represented in a numerical form. As it follows from (6.149), when using the calculated values of Hs1 and Hs2 determined by the expressions (6",
            "154) will contain only two independent constants, since for their constants, in this case, the conditions of the form z11\u00bc z22 and z12\u00bc z21 are satisfied. In this regard the constants of the system of Eq. (6.154) can be expressed through the known numerical values of Es1,Hs1 and Es2,Hs2 given on the outer and inner surfaces of the sth layer. As a result, we have z11 \u00bc z22 \u00bc Es1Hs1 \u00fe Es2Hs2 H2 s1 H2 s2 ; z12 \u00bc z21 \u00bc Es1Hs2 \u00fe Es2Hs1 H2 s1 H2 s2 \u00f06:155\u00de The system of Eq. (6.154) can be used to construct the equivalent circuit for the sth layer of the model in Fig. 5.1b. As was shown above, the reactance values used in the equivalent circuit of the sth layer are expressed through the |z| \u2013 constants of the system of Eq. (6.154). Therefore with the use of expression (6.155), these reactance values can be represented through the numerical values of Es1,Hs1 and Es2,Hs2 determined on the outer and inner surfaces of the sth layer. Magnetizing Reactance Values of the Air Gap. Expressions for the air gap magnetizing reactance values represented through the |z| \u2013 constants are given in (6",
            " When the squirrel-cage type winding is used in the rotor slots, the layer of the wound part of the rotor tooth region is furthermore characterized by the average value of the resistivity determined by the expression given in (3.51) and (7.45) in the following form \u03c1zR \u00bc \u03c12 tzR b\u03a0R \u00f08:1\u00de Therefore, in this case the wound part of the rotor tooth region is represented as the conducting magneto-anisotropic layer. 276 8 Single-Winding Rotor Induction Machine Circuit Loops: Weak Skin Effect In accordance with Fig. 5.1, the stator and rotor of an electric machine are subdivided into regions (layers). While under load, currents flow in the layers reflecting the wound parts of the stator and rotor teeth regions (layers zs and zR in Fig. 5.1). Currents in the rotor layer zR are induced (when using a closed phase or squirrel-cage type windings). In the stator layer zs, the current is caused by voltage applied to the terminals of the stator winding. The phase stator and rotor windings are represented as infinitely thin filaments of the currents distributed uniformly in the wound parts of their teeth regions. Therefore, in the model in Fig. 5.1, the stator layer zs represents the active layer. When using a phase winding on the rotor, the current in it can be caused by an external source. Then, in the model in Fig. 5.1, the wound part of the rotor tooth region is considered as the active layer. The resistance of the phase winding is represented as the external resistance. In order to determine this resistance the active part of the phase winding is considered as the conducting conditional bar located in the slots. In accordance with the expressions given in (3.42) and (3.46), the resistivity of this conditional bar is determined as \u03c1 0 k \u00bc (\u03c1k/k\u0417kk2wk), where \u03c1k is the resistivity of the winding material, k\u0417k is the fill factor of the slot, and k\u00bc 1, 2. Next, this conducting conditional bar is represented as a conducting layer formed by the cooper of the winding. Based on expressions given in (3.43) and (3.47), the average resistivity of this conditional layer takes \u03c1zi \u00bc \u03c1 0 k tzi b\u03a0i \u00f08:2\u00de where k \u00bc 1, 2; i \u00bc s,R So, in accordance with (8.2), determining the resistance of the phase winding is implemented on the basis of the representation of the phase winding as the conditional conducting layer. According to Fig. 5.1, in relation to the surface passing through the middle of the air gap, the electric machine model breaks down into a two-layer system: the inner layers represent the rotor model and the outer layers reflect the stator model. In this regard, the construction process of an electric machine equivalent circuit can be divided into three phases. In the first phase, an equivalent rotor circuit is built. In the second phase, the equivalent stator circuit is constructed. An equivalent electric machine circuit arises due to the cascade connections of the equivalent stator and rotor circuits",
            " As a result, a multi-loop circuit with a mixed connection of elements arises for an electric machine. Then, this equivalent circuit can be converted into a multi-loop circuit with parallel element connections. A multi-loop circuit with parallel connection of the elements represents the circuit loops of an induction machine. Features of the construction of equivalent circuits and determining the circuit loops for an induction machine with the single-winding rotor are considered below based on the model shown in Fig. 5.1. This model allows description of the resulting field in the active zone of an electric machine by the Maxwell\u2019s equations. 8.1 Induction Machine with a Single-Winding Rotor: Multi-Layer Model 277 In accordance with the model in Fig. 5.1, the rotor of an electric machine is represented consisting of four layers: the air gap sub-layer (with a length equal to \u03b4/2 adjacent to the rotor surface); layers of the slot wedge and wound part of the tooth region; and a joke region layer. In Chap. 5 the modular method for construction of an equivalent circuit for the multi-layer structure was described. We use this method for constructing equivalent circuits of the single-cage rotor. In accordance with the modular method, it is necessary to first construct equivalent circuits for the rotor model layers",
            " In technical literature these expressions arise on the basis of direct consideration of the leakage field in the rotor slot, in which the winding bars with the electromagnetic parameters \u03c12 and \u03bc0 are located. 0 0 Em HR x02 x\u03c402 x\u03c4kR s rc2 x\u03a02\u03c3 \u00b7 \u00b7 Fig. 8.10 Single circuit loop representation of a single-cage rotor for the weak skin effect 8.2 Single-Cage Rotor: Circuit Loops 297 The conductors of the phase winding in the rotor slots are withheld by the slot wedges. While under load the rotor phase winding is usually closed on some external resistance. Generally, we believe that the rotor winding is supplied from an external source (exciter). Therefore, the wound part of the rotor tooth region in Fig. 5.1 is represented as the active non-conducting magneto-anisotropic layer. The electromagnetic parameters \u03bcxzR, \u03bcyzR and \u03c1zR of this active layer can be determined by the expressions given in (3.43), (3.47), (5.1), (5.2), (5.3) and (8.2). An equivalent circuit of a phase winding rotor can be constructed using the modular method described above. Here, to simplify construction of such equivalent rotor circuits we use the single-cage rotor equivalent circuits obtained in Figs. 8.2 and 8.3. For this purpose, the fragments representing the wound part of the rotor tooth region in these equivalent circuits can be replaced by the circuits reflecting the currents in the rotor phase winding. In the model in Fig. 5.1, the wound part of the rotor tooth region is represented as the active layer. The equivalent circuits for the active layer were obtained in Chap. 7 for the general case. In Fig. 7.9 the T-circuit is shown and in Fig. 7.11 the L-circuit has been given for the active layer. In our case, the active rotor layer represents the non-conducting magneto-anisotropic medium. Therefore, the equivalent circuit elements obtained in Figs. 7.9 and 7.11 acquire an inductive character. As applied to our case the equivalent circuits in Figs",
            "17, the four parallel branches with the reactance values x02, xkR, xzR and x\u03c4azR can be united into a single common branch, the reactance of which is determined as 308 8 Single-Winding Rotor Induction Machine Circuit Loops: Weak Skin Effect xmR \u00bc 1 1 x02 \u00fe 1 xkR \u00fe 1 xzR \u00fe 1 x\u03c4azR \u00bc x02 1 1\u00fe x02 1 xkR \u00fe 1 xzR \u00fe 1 x\u03c4azR \u00bc x02 1 kHR \u00f08:91\u00de where kHR \u00bc 1\u00fe x02 1 xkR \u00fe 1 xzR \u00fe 1 x\u03c4azR This expression is completely consistent with the formula (8.49) for the magnetizing reactance of the single-cage rotor. On the basis of the expression (8.91), the equivalent circuit of the phase winding rotor takes the form in Fig. 8.18. In this equivalent circuit, xmR represents the phase winding rotor magnetizing reactance. The equivalent circuit in Fig. 8.18 reflects the rotor phase winding circuit loop. In accordance with the model in Fig. 5.1, the stator is represented as four layers: the air gap sub-layer with a length equal to \u03b4/2 (adjoining to the stator bore surface); the layers of the slot wedge and wound part of the tooth regions; and the stator joke region layer. According to the method described above, construction of the stator equivalent circuit is implemented in two stages. In the first stage, the equivalent circuits for the stator layers and air gap sub-layer are constructed. Then, in the second stage these equivalent circuits are connected with each other in a cascading fashion taking into account the spatial arrangement of the layers in the stator model (Fig. 5.1b). As a result, a ladder circuit or multi-loop circuit with mixed element connections arises for the stator. Then the stator ladder circuit can be converted into a multi-loop circuit with parallel element connections. Based on this circuit, the stator circuit loops can be established. We realize this provision below. The layers of the stator model (Fig. 5.1b) can be replaced by T- or L-circuits. When using T-circuits for the stator model layers, the stator equivalent circuit arises as a result of the T-circuits cascade connection. In this case, we can use the equivalent circuit given in Fig. 5.4 for the air gap sub-layer adjoining the stator bore surface. On the basis of Fig. 7.3, the T-circuit for the stator slot wedge region 8.4 Single-Winding Stator: Circuit Loops 309 takes the form in Fig. 8.19. Reactance values of this equivalent circuit can be determined by the expressions (7",
            " Then, it follows xks0 \u00bc \u03c91\u03bcyks \u03b2kssh\u03b2kshks \u03beLks \u03be\u03c4ks\u03b6ks ; x\u03c4ks1 \u00bc x\u03c4ks2 \u00bc \u03c91\u03bcxks \u03b2ks \u03beLks \u03be\u03c4ks ch\u03b2kshks 1 sh\u03b2kshks \u03b6\u03c4ks \u00f08:92\u00de where \u03b2ks \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi k\u03bcks\u03c0=\u03c4ks q ; \u03be\u03c4ks \u00bc \u03c4ks=\u03c40; \u03beLks \u00bc lks=l0 Here, the curvature factors \u03b6ks and \u03b6\u03c4ks represent the ratio of the reactance values xks0 and x\u03c4ks1\u00bc x\u03c4ks2 determined in cylindrical and planar coordinate systems. Taking into account the first two terms of a series expansion of the hyperbolic functions, we can receive from (8.92) for the reactance values xks0 and x\u03c4ks1\u00bc x\u03c4ks2 xks0 \u00bc \u03c91\u03bcyks\u03c4 2 ks \u03c02hks \u03beLks \u03be\u03c4ks\u03b6ks ; x\u03c4ks1 \u00bc x\u03c4ks2 \u00bc \u03c91\u03bc0tks hks 2b\u03a0s \u03beLks \u03be\u03c4ks \u03b6\u03c4ks \u00f08:93\u00de The stator winding is supplied from the external source. Therefore, the wound part of the stator tooth region is represented as the active layer in the stator model in Fig. 5.1b. The \u0422-circuit of this active layer arises from the equivalent circuit shown in Fig. 7.9. The stator active layer is represented as the non-conducting magnetoanisotropic medium. Therefore, the elements of the equivalent circuit in Fig. 7.9 acquire, in this case, an inductive character. In the equivalent circuit in Fig. 7.9, index 2 can be replaced by index 1. In addition, we can assume that i\u00bc zs. Then, the equivalent circuit in Fig. 7.9 is converted into the equivalent circuit in Fig. 8.20. The equivalent circuit in Fig",
            "4 Single-Winding Stator: Circuit Loops 311 xzs0 \u00bc \u03c91\u03bcyzs\u03c4 2 zs \u03c02h\u03a0s \u03beLzs \u03be\u03c4zs\u03b6zs ; x\u03c4zs1 \u00bc x\u03c4zs2 \u00bc \u03c91\u03bc0tzs h\u03a0s 2b\u03a0s \u03beLzs \u03be\u03c4zs \u03b6\u03c4zs x10 \u00bc \u03c91\u03bc0tzs h\u03a0s 6b\u03a0s \u03beLzs \u03be\u03c4zs \u00f08:97\u00de Now connecting in a cascading manner the equivalent circuits given in Fig. 5.4 for the air gap sub-layer, in Fig. 8.19 for the slot wedge layer, in Fig. 8.20 for the layer of the wound part of the stator tooth region, and also considering that the stator joke layer is replaced by a two-terminal network (Fig. 5.11), the stator equivalent circuit takes the form in Fig. 8.21. The stator equivalent circuit in Fig. 8.21 was obtained using T-circuits. The layers of the stator model (Fig. 5.1b) can be replaced by the L-circuits. The equivalent stator circuit arises, in this case, as a result of the L-circuit cascade connection. For the air gap sub-layer, we have the equivalent circuit given in Fig. 5.4. The expressions for the reactance values of the equivalent circuit given in Fig. 5.4 were obtained in (6.79) and (6.84). The stator slot wedge layer is replaced by the L-circuit in Fig. 5.8. The reactance values of the L-circuit in Fig. 5.8 are defined by the expressions given in (6.66), (6"
        ],
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    },
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        "image_filename": "designv11_13_0003915_j.jmapro.2019.05.002-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003915_j.jmapro.2019.05.002-Figure5-1.png",
        "caption": "Fig. 5. Model Mesh and Elements activation.",
        "texts": [
            " Since the major heat source in the process is the laser beam and the heat generated during the mechanical deformation is negligible, an uncoupled model, as done by other researchers [12], is considered. In this model, the temperature distribution is calculated in the first step to be used for the mechanical model. For modeling the clad formation, solid elements attached to the original substrate are initially defined in the model and deactivated at the start of the process in the FE software. During the simulation of the process, these elements are activated in a step-wise manner every 0.33 s so that 39 steps, in 13 s, are used to create each bead. The Fig. 5 shows the details of the FE modeling. In the following, the basics of the thermal and mechanical analyses are described. The governing equation of the transient heat conduction is: \u2202 \u2202 = \u2212 \u2207 +\u03c1C T t k T Qp 2 (1) where \u03c1 is the density, Cp is the specific heat capacity, k is the thermal conductivity, T is the temperature, and Q is the power of heat source per unit volume. In the present analysis, Q is produced by laser radiation. The associated boundary conditions are as follows: \u2212 \u2207 = \u23a7 \u23a8\u23a9 \u2212 \u2212 \u2212 \u2212 \u2208 \u2212 \u2212 \u2212 \u2212 \u2209 k T \u03b1D x y z t h T T \u03b5 \u03c3 T T \u03c8 \u0393 h T T \u03b5 \u03c3 T T \u03c8 \u0393 ( "
        ],
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    },
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        "image_filename": "designv11_13_0000330_1.4029776-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000330_1.4029776-Figure2-1.png",
        "caption": "Fig. 2 Schematic of noncontacting finger seal",
        "texts": [
            " The nature of dynamic leakage performance affected by the change of seal\u2013rotor clearance is revealed. There are three steps for dynamic leakage analysis. In the first step, closed-form description of seal\u2013rotor clearances is obtained through dynamic model [12,15]. In the second step, mass flow rate under finger mass is obtained. In the third step, seal leakage analysis is coupled to dynamic analysis to get mass flow rate of a full seal. The detailed flow diagram of the proposed method is illustrated in Fig. 1. The two-dimensional schematic of the noncontacting finger seal is shown in Fig. 2 [3,4]. The purpose of back plate is to serve an axial support against high pressure (HP) and low pressure (LP) fingers\u2019 bending. The LP finger has a hydrodynamic and hydrostatic self-acting lifting pad. The LP pad generates a small sealing clearance between rotating shaft and finger pad to provide selfacting lifting force. A Cartesian coordinate system is used for rotor because the thickness of gas film thickness underneath LP finger pad is much smaller than seal radius, and curvature of gas film is omitted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001009_iet-epa.2016.0680-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001009_iet-epa.2016.0680-Figure10-1.png",
        "caption": "Fig. 10 Test bench of 1.1 kw IM",
        "texts": [
            " The above results are consistency with derived (37) and (48). Simulation results of fixed flux higher than rated value at different reference torques are illustrated in Fig. 9. As can be seen from this figure, the rule described by (37) and (48) is still hold in this condition. The errors between values calculated by (27), (39), (40) and actual values are within 3%, which shows the good accuracy of derived formulas. Experimental tests were carried out to verify the correctness of derived conclusions. The test bench, shown in Fig. 10, is made up of a 1.1 kw ABB IM with an Omron 2000 pulses optical encoder, a 10Nm torque sensor, a digital signal processor (DSP) controller and a 1.3 kw Yasukawa servo motor. The controller is composed of a Texas Instruments (TI) DSP TMS320F28069 control card, an insulated gate bipolar transistor VSI, a 220 V AC to 350 V DC rectifier, three-phase current and voltage sample circuit, controller area network (CAN) communication circuit and protection circuit. The whole control scheme is built by Simulink model, which can generate C code automatically using real-time workspace of MATLAB"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003193_j.aca.2017.09.027-Figure587-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003193_j.aca.2017.09.027-Figure587-1.png",
        "caption": "Figures 587",
        "texts": [],
        "surrounding_texts": [
            "The principle of c-EME is illustrated in Figure 1a. The supported liquid membrane (SLM) is a key 96 element in the set-up. The SLM consists of < 1 \u00b5L of an organic solvent, which has been 97 immobilized in the pores of a small piece of porous hollow fiber (or in a flat porous membrane as 98 discussed below). The organic solvent is non-polar and immiscible with water. The piece of 99 hollow fiber with the SLM is located in a flow path of acceptor solution. During EME the SLM is in 100 contact with the sample. In this tutorial, the sample is an in-vitro drug metabolism solution, but 101 c-EME is not limited to this application. The sample can be stagnant (in a fixed position, but 102 stirred) as illustrated in Figure 1a or the sample can be dynamic in a flow system as illustrated 103 in Figure 1b+c. 104 As illustrated in Figure 1a, target compounds are extracted continuously from the sample, 105 through the SLM and into the acceptor solution. The driving force for extraction is an electrical 106 field (extraction voltage), which is sustained across the SLM. For extraction of cationic 107 compounds, the cathode (negative electrode) is located in the acceptor solution. For extraction 108 of anionic compounds, the polarity of the electrical field is reversed. Neutral species are not 109 extracted by the applied voltage. The polarity and magnitude of the electrical field is controlled 110 from an external power supply. The acceptor solution is continuously pumped by a micro-syringe 111 pump, and in this way the extracted analytes are transferred into a UV-spectrophotometer or a 112 mass spectrometer for continuous on-line detection. One practical example is illustrated in 113 Figure 2, where continuous EME was coupled directly and on-line to mass spectrometry for fast 114 and real-time characterization of in-vitro metabolism of the antidepressant drug amitriptyline. 115 The concentration of amitriptyline in the reaction mixture was followed with the mass 116 spectrometer, which measured the protonated molecular ion (MH+) for amitriptyline at m/z 117 278. Thus, the m/z 278 trace shows the time-dependent metabolism of amitriptyline, which was 118 initiated after 13 minutes at point c) in Figure 2. Also, the mass spectrometer measured several 119 metabolites formed during the experiment, namely hydroxy-amitriptyline at m/z 294, 120 dihydroxy-amitriptyline at m/z 310, and nortriptyline at m/z 264. More details about this 121 experiment are discussed later. Traditionally, metabolic data similar to those in Figure 2 are 122 acquired by pipetting and collection of reaction mixture at certain time-points, followed by 123 protein precipitation by addition of example acetonitrile to each sample to terminate the 124 reaction. Finally, all the samples are analyzed sequentially by LC-MS. With c-EME coupled 125 directly to mass spectrometry, the entire experiment was completed and all data were acquired 126 in 45 minutes. In this case, c-EME served as sample preparation unit in front of the mass 127 spectrometer. Thus, the main purpose of c-EME in this experiment was to isolate the drug 128 substances (parent drug and metabolites) from buffer components, salts, and proteins present 129 in the in-vitro reaction mixture as well as to stop the ongoing metabolism in the extract. The 130 reaction mixture was too complex for direct introduction into the mass spectrometer. 131 3. Technical and operational challenges \u2013 part I 132 Ideally, c-EME with on-line detection should provide reliable and time-resolved data with 133 minimal time delay. Thus, compositional changes in the sample should immediately be 134 measured correctly by the detector. In addition, the system should be capable of measuring 135 target compounds at low concentration levels. These requirements introduce the following 136 general challenges related to design and operation: 137 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 5 \u2022 Extraction recovery should be constant over the entire time frame of an experiment - 138 criterion (1) 139 \u2022 Transfer of target compounds from sample and to the SLM should be very fast - criterion 140 (2) 141 \u2022 Transfer of target compounds across the SLM and into the acceptor solution should be 142 very fast - criterion (3) 143 \u2022 Post-EME transfer of compounds to on-line detection should preferably be very fast - 144 criterion (4) 145 \u2022 Extraction recovery should be high - criterion (5) 146 When c-EME is used to study in-vitro metabolic reactions, the following additional challenges 147 are added to the list: 148 \u2022 The temperature of the reaction mixture should be 37\u02daC - criterion (6) 149 \u2022 The reaction mixture should have sufficient access to open air, both for supply of oxygen 150 and as well to easily addition of reagents to the mixture - criterion (7) 151 Criterion (1) is important to ensure reliable quantitative data, because highly stable conditions 152 are a prerequisite for consistent recoveries, which in turn are vital for accurate correlation 153 between concentration in the sample and the measured detector signal. Criterion (1) requires a 154 stable SLM and constant pH conditions in sample and acceptor solution. Criteria (2), (3), and (4) 155 are important to maintain a fast response in the system. Fast extraction kinetics is very 156 important to prevent that the acquisition of drug metabolism kinetic profiles is influenced by 157 the extraction process. Criterion (5) is important in order to maintain high sensitivity and to 158 enable measurements at low concentrations. Criterion (5) is not valid for soft extraction as 159 discussed later. Criteria (6) and (7) are important in order to acquire metabolic data under 160 physiologically relevant conditions. The technical and operational solutions to address the 161 criteria are discussed in the following, and often the criteria are challenging to fulfil. 162 4. Theory 163 Several equations have been developed to describe extraction recovery in EME/c-EME [12,18]. 164 Although all the equations are based on assumptions, they can be used to identify the main 165 operational parameters. Under both stagnant and dynamic sample conditions, high distribution 166 of the target compound into the SLM is crucial in order to obtain fast extraction across the SLM 167 (criterion (3)). Also, high distribution is a prerequisite for high extraction recovery (criterion 168 (5)). The distribution (Kd*) can be described by the following equation [19]: 169 ( )      \u2206\u03a6\u2212\u2206\u03a6= 0* exp i i d RT Fz K (1) 170 \u0394\u03a6 is the Galvani potential difference across the sample-SLM interface, and this is proportional 171 to the extraction voltage. \u0394\u03a6i 0 is a target compound dependent standard extraction potential 172 related to the hydrophobicity, zi is the charge of the target compound, F is Faradays constant, R 173 is the gas constant, and T is the absolute temperature. From equation (1) it is clear that the 174 extraction efficiency of a given substance is dependent on: 175 \u2022 Extraction voltage (operational parameter) 176 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 6 \u2022 Target compound hydrophobicity (compound-related parameter) 177 \u2022 Temperature 178 Thus, the extraction voltage is an important operational parameter, and generally extraction 179 kinetics and recoveries improve with increasing voltage up to a certain level. Above this voltage, 180 the rate limiting step is not transfer across the SLM, and equation (1) is no longer valid. Also, 181 kinetics and recoveries are related to the hydrophobicity of the target compound (compound-182 related parameter), and both are improved with increasing hydrophobicity up to a certain level. 183 Temperature is also affecting the extraction, but this is normally kept constant (37 \u0366C in case of 184 in-vitro drug metabolism) and is not considered an operational parameter. 185 Under stagnant sample conditions using the probe configuration, the time-dependent extraction 186 of target compounds from the sample can be described by the following equation [19]: 187 ( )       \u22c5 \u22c5\u2212 \u22c5= \u2192 t V PA CtC D ADf DD exp0 (2) 188 CD(t) is the time-dependent concentration of target compound in the sample; CD 0 is the original 189 concentration of target compound in the sample; Af is the active surface of the SLM; PD\u2192A is a 190 compound dependent SLM permeability coefficient (related to Kd*); VD is the sample volume; 191 and t is time. From equation (2) it is clear that the extraction efficiency of a given target 192 substance from stagnant sample is dependent on the following parameters in addition to 193 extraction voltage (equation (1)): 194 \u2022 Active SLM surface (device related parameter) 195 \u2022 SLM permeability coefficient (compound-related parameter, partially linked to Kd*) 196 \u2022 Sample volume (operational parameter) 197 \u2022 Time (operational parameter) 198 The active SLM surface is an important parameter to consider during construction of devices, 199 and increasing the surface area increases the efficiency of EME. Sample volume and time are 200 important operational parameters, and increasing time while decreasing sample volume is both 201 in favour of high efficiency. 202 Under dynamic sample conditions, the extraction recovery R can be described by the following 203 equation [12]: 204 ( ) %1001 /0 \u22c5\u2212= \u2212 svameR (3) 205 The term a is the sample contact area with the SLM; m0 is a compound dependent mass transfer 206 coefficient; and vs is the volumetric flow rate of sample. From equation (3), it is clear that the 207 extraction efficiency of a given target substance under dynamic sample conditions is dependent 208 on the following parameters in addition to extraction voltage (equation (1)): 209 \u2022 Active SLM surface (device related parameter) 210 \u2022 Mass transfer coefficient (compound-related parameter, partially linked to Kd*) 211 \u2022 Sample flow rate (operational parameter) 212 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 7 Thus, under dynamic sample conditions, the sample flow rate is an important operational 213 parameter, and extraction efficiency is increasing with decreasing flow rate. 214 In addition to the operational parameters discussed above, also the ion balance (\u03c7) across the 215 SLM can impact extraction performance [20]. The ion balance is the ratio of the total ionic 216 concentration on the sample side to that on the acceptor solution side. Theoretically, a decrease 217 of the ion balance results in increased flux of target compound across the SLM. This has been 218 verified and observed experimentally in conventional EME systems [20], but no such 219 experiments have currently been reported for c-EME. Therefore, ion balance experiments 220 should be included in future c-EME work. 221 5. Micro-chip EME, technical design 222 As discussed above, we have performed c-EME in three different technical configurations up to 223 date, namely (a) micro-chip, (b) probe, and (c) flow-probe configuration. In the following, we 224 discuss the micro-chip configuration, which is illustrated in Figure 3. This EME-chip has been 225 presented in several research papers [11-13]. The EME-chip was composed of two polymethyl 226 methacrylate (PMMA) plates, and channels for sample and acceptor solution were milled in the 227 plates. The channels were 6 mm long and rectangular, with a depth of 50 \u00b5m and a width of 2.00 228 mm. The active volumes of sample and acceptor solution were approximately 0.6 \u00b5L (600 nL) 229 each, and diffusion distances to the SLM were very short. The latter was considered important 230 for extraction kinetics (criterion (2)) and recovery (criterion (5)). The active surface area of the 231 SLM was 12 mm2. The two plates were clamped, with a piece of commercially available porous 232 polypropylene membrane in between. This membrane served as support for the SLM, and had 233 55 % porosity and 0.5 x 0.05 \u00b5m2 pores. The thickness of the membrane was 25 \u00b5m, and 234 therefore the SLMs were very thin in micro-chip c-EME as compared to traditional EME (100-235 200 \u00b5m). A thin membrane was selected to ensure fast extraction across the SLM (criterion (3)). 236 The micro-chip was assembled and fixed by solvent-assisted bonding and the channels were 237 equipped with connecting tubes as described in details in the original papers [11-13]. 238 Coupling of the electrical field is crucial in EME, and ideally both electrodes should be located 239 very close to the SLM. The reason for this is to ensure that the voltage drop mainly occurs across 240 the SLM for fast extraction (criterion (3)). In principle, the electrodes can be located far from the 241 SLM everywhere in the flow path of sample and acceptor solution, but in such cases (a) a large 242 portion of the voltage drop occurs in the aqueous solutions, (b) exact control of the electrical 243 field across the SLM becomes more challenging, and (c) high voltages have to be used. In the 244 micro-chip illustrated in Figure 3, we solved this important issue by using a stainless steel 245 tubing as inlet for acceptor solution and coupling this to the anode outlet of the external power 246 supply. Thus, the distance between the anode and the SLM was only a few millimeters. The 247 cathode comprised in a similar way a piece of stainless steel tubing, which was inserted in the 248 sample flow path a few centimeters from the SLM. The cathode was located in the sample outlet. 249 In this way, any potential bubble formation, pH changes or other electrode-related issues were 250 not influencing the extraction process. 251 The sample chamber was coupled directly to the micro-chip as illustrated in Figure 3. In this 252 case, the sample chamber was used for study of in-vitro metabolism of drug substances. Such 253 experiments are conducted at physiological conditions, which imply pH 7.4 and 37\u02daC in the 254 reaction mixture. Appropriate pH conditions were maintained using a phosphate buffer, while 255 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 8 the temperature was maintained by thermostatic control using circulating water (criterion (6)). 256 Oxygen plays an important role in drug metabolism, and access to air is mandatory for the 257 metabolic experiments. Therefore, the sample chamber was open to air, and sample was sucked 258 into the EME-chip using a micro-syringe pump at the sample outlet (criterion (7)). The micro-259 syringe also served as the final sample waste reservoir, and drug substances and biological 260 waste were confined to the micro-syringe for safe disposal. The distance from the sample 261 chamber to the EME-chip was only a few centimeters, to minimize the transfer time (criterion 262 (2)). 263 The acceptor solution was pumped into the EME-chip using a second micro-syringe pump. 264 Because the membrane supporting the SLM was flexible, the shape of the membrane was 265 affected by the internal acceptor solution pressure, which in turn was determined by the flow-266 rate of the acceptor solution (at a given flow-rate of sample). Therefore, flow-rates were 267 generally not exceeding 20 \u00b5L min-1. After EME, the analytes were transferred to the mass 268 spectrometer for continuous on-line detection. The acceptor outlet of the micro-chip was 269 connected to the electrospray ionization source of the mass spectrometer by a fused silica 270 capillary. Again, to minimize the transfer time, the capillary was kept as short as possible and 271 therefore the micro-chip was located as close as possible to the mass spectrometer (criterion 272 (4)). 273 6. Micro-chip EME, operational parameters and performance 274 Typical operational parameters for micro-chip c-EME are shown in Table 1 when we used the 275 system in combination with mass spectrometry for in-vitro drug metabolism studies. The 276 antidepressant drug amitriptyline was used as model substance in these experiments. 2-277 Nitrophenyl octyl ether (NPOE) was used as SLM. Due to the high hydrogen bond basicity of 278 NPOE and due to the high log P, this solvent has been identified as a very efficient solvent for 279 EME of protonated non-polar basic drugs (log P > 2) [10]. NPOE has by far been the most 280 popular solvent for traditional EME of basic substances. In c-EME however, the stability of the 281 SLM becomes much more critical because the SLM is used for long periods of time, and because 282 the SLM is operated under flow conditions. In spite of this, we have experienced that NPOE is 283 highly stable as SLM even in c-EME (criterion (1)). The sample was a metabolic reaction 284 mixture, comprising (a) rat liver microsomes, (b) phosphate buffer, (c) MgCl2, (d) cofactor 285 (NADPH), and (e) the drug substance to be examined [13]. The sample was sucked into the 286 EME-chip with a flow rate of 20 \u00b5L min-1. This relatively high flow rate was not optimal in 287 relation to extraction recovery (equation (3)), but provided the highest signals for amitriptyline 288 in the mass spectrometer. A 10 mM solution of formic acid was used as acceptor solution. Acidic 289 conditions were mandatory to keep the drug substances protonated. This was important for 290 efficient c-EME and for efficient mass spectrometric detection. In the range up to 20 \u00b5L min-1, 291 the extraction efficiency was relatively insensitive to the flow rate of acceptor solution. 292 However, both the mass spectrometric signals and the transfer time from the micro-chip were 293 improved at higher flow-rates and therefore 20 \u00b5L min-1 was selected. The extraction voltage 294 was set to 15 V. With our EME-chip systems, the extraction efficiency increased with increasing 295 voltage up to 10-15 V according to equation (1), but above this level there was no further gain in 296 efficiency. 297 Under the experimental conditions reported in Table 1, amitriptyline was extracted with a 298 recovery of 48 % from the flowing sample [13]. This value was increased to 83 % when the 299 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 9 sample flow rate was reduced from 20 to 5 \u00b5L min-1 [13], and this was in accordance with 300 equation (3). With reduced sample flow rate, the residence time in the micro-chip increased. In 301 another paper, we tested extraction where the flow of acceptor solution was stopped (stop-flow 302 mode) [12]. With methadone as target compound, enrichment by a factor of 27 was obtained 303 after 4 minutes of operation, and by a factor of 76 after 12 minutes of EME. After extraction in 304 stop-flow mode, the acceptor solution flow was turned on again to transfer the extracted 305 compounds to the mass spectrometer. 306 Our micro-chip EME systems have been working very efficiently in the lab [11-13]. Performance 307 data have been at the level reported above, and repeatability as well as linearity has been 308 acceptable. When we first learned how to address the challenges discussed above and how to 309 best operate the system, the system was operated more or less continuously without major 310 failure or technical issues. We think micro-chip EME shows great potential for applications in 311 the future, but one important challenge requires additional work: The porous membrane 312 support for immobilization of the SLM is located inside the chip. Therefore, to load the SLM, the 313 connecting tubes have to be removed and the SLM solvent has to be pipetted into the membrane 314 support. Excess organic solvent is then flushed from the micro-chip, and this may contaminate 315 channels and tubing. Perhaps the most elegant way to solve this issue is to develop single-use 316 micro-chips, which are disposed after one extraction, and where the SLM is loaded during 317 fabrication of the micro-chip. 318 7. Probe EME, technical design, operational parameters and 319 performance 320 c-EME has also been reported in probe and flow-probe configurations as illustrated in Figure 321 4a+b. In the probe configuration (Figure 4a), extractions were conducted from stagnant sample 322 [14]. A small segment of porous polypropylene hollow fiber (1.0-5.0 mm active length, 330 \u00b5m 323 internal diameter, and 70 \u00b5m wall thickness) was used as support for the SLM. The active 324 volume of acceptor solution was in the range 90-430 nL depending on the length of the hollow 325 fiber, and the corresponding active surface area of the SLM was in the range 2-10 mm2. The inlet 326 of the hollow fiber segment was heat-sealed to a steel capillary, and this was connected to a 327 micro-syringe pump for delivery of acceptor solution. The steel capillary also served as cathode. 328 The outlet of the hollow fiber segment was heat-sealed to a fused silica capillary, and this was 329 connected to the mass spectrometer. The anode (platinum wire) was inserted in the reaction 330 chamber, and the latter was temperature controlled at 37 \u030aC. Applicability of this system was 331 exemplified by studies of in-vitro metabolic reactions using amitriptyline as model drug under 332 the following experimental conditions: The sample volume was 1000 \u00b5L (metabolic reaction 333 mixture), NPOE (< 1 \u00b5L) was used as SLM, 3 \u00b5L min-1 of 60 mM formic acid was used as acceptor 334 solution, and the extraction voltage was set to 2.5 V. With this low voltage, the extraction 335 efficiency was reduced and the system was tuned for soft extraction. Soft extraction was 336 performed in this case in order not to affect the kinetics of the metabolic reaction by c-EME, 337 since the extraction was conducted directly from the bulk reaction mixture. The probe-EME 338 configuration has several pros: (a) the probe was mechanically very flexible and technically very 339 simple, (b) the probe was well suited for in-vitro drug metabolism studies and provided reliable 340 data, and (c) loading the SLM was very convenient since the membrane support was accessible 341 at the end of the probe. The major cons with probe-EME are related to the fact that extraction 342 removes substances of interest from the sample (reaction mixture), and this may affect 343 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 10 chemical/biochemical equilibria. In such cases, the probe-EME system has to be operated under 344 soft extraction conditions, but this is at the expense of sensitivity. 345 To address the latter challenge, a flow-probe was developed for c-EME as illustrated in Figure 346 4b [15,16]. In this system, small portions of the sample (reaction mixture) were sucked 347 continuously into a piece of PTFE tubing by a micro-syringe pump, and inside this tubing c-EME 348 was conducted. In this way, extraction was performed on a small portion of sample which was 349 physically separated from the bulk sample. Therefore, although the total volume of sample 350 gradually decreased during extraction, the composition within the bulk sample was not affected 351 by c-EME. The hollow fiber support, SLM, and flow path for acceptor solution were very similar 352 to the probe discussed above. The use of highly flexible fused silica capillaries for the acceptor 353 solution enabled the flow-probe construction as shown in Figure 4b, and the capillary 354 connecting the flow-probe to the mass spectrometer was extended through the wall of the PTFE 355 tubing. This arrangement was feasible and leak tight due to the soft and flexible nature of the 356 PTFE tubing. The applicability of flow-probe-EME was exemplified by studies of in-vitro 357 metabolism using promethazine and methadone as model drugs, under the following 358 conditions: The sample volume was 2000 \u00b5L (metabolic reaction mixture), the removal rate of 359 sample was 10 \u00b5L min-1, NPOE (< 1 \u00b5L) was used as SLM, 10 \u00b5L min-1 of 10 mM formic acid was 360 used as acceptor solution, and the extraction voltage was 200 V. A high extraction voltage was 361 used with the flow-probe because c-EME was not performed directly from the bulk reaction 362 mixture. Operation at high voltage was associated with two important advantages; (a) the 363 sensitivity increased due to increased extraction recovery and (b) extraction kinetics were 364 improved (both according to equation (1)). Rapid extraction kinetics is mandatory in order to 365 measure correct kinetic data for the drug metabolism reaction. Concentration versus time 366 profiles obtained with the EME flow-probe coupled directly to mass spectrometry were 367 comparable with profiles obtained by traditional sampling at defined time points followed by 368 protein precipitation and LC-MS [15]. Thus, the flow-probe EME system provided reliable 369 kinetic information. 370 Very recently, the c-EME flow-probe has been advanced further by introducing a make-up flow, 371 which was mixed with the sample flow prior to c-EME [17]. This enabled c-EME to be 372 accomplished at other pH-values than 7.4, which is the pH of the in-vitro metabolic reaction 373 mixture. With hydrochloric acid as make-up flow, samples were acidified for extraction of 374 zwitterionic drug metabolites of vortioxetine and hydroxyzine [17]. 375 8. Applications 376 In the work reported up to date, the applicability of each systems for c-EME has been 377 exemplified by in-vitro drug metabolism using different non-polar basic drugs as model 378 substances. One example is shown in Figure 2 with amitriptyline (m/z 278 for MH+). At the on-379 set of the experiment (t = 0.0 min), the reaction chamber only contained phosphate buffer pH 380 7.4 and dissolved MgCl2. Thus, no amitriptyline was present in the reaction mixture and only 381 background signal was measured at m/z 278 with the mass spectrometer. After 4.0 minutes of 382 experiment (a), amitriptyline was added to the reaction mixture, and shortly thereafter the 383 signal at m/z increased accordingly. At 5.0 minutes, the m/z 278 signal was stabilized, and this 384 signal represented the total concentration of amitriptyline in the reaction mixture (10 \u00b5M). At 385 7.3 minutes (b), rat liver microsomes (enzymatic proteins) were added to the reaction mixture. 386 The signal at m/z 278 decreased and then stabilized, due to the dilution effect of the rat liver 387 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 11 microsome suspension and probably also partly due to some protein binding of amitriptyline. 388 After a short time, the m/z 278 signal stabilized again, and at 12.7 minutes (c) cofactor (NADPH) 389 was finally added to the reaction mixture. This initiated the metabolic reaction, and the signal at 390 m/z 278 decreased now as function of time as shown in Figure 2. The logarithmic decrease in 391 concentration supported first-order kinetics, and the half time for amitriptyline was calculated 392 to 1.4 minutes under the given experimental conditions. The mass spectrometer was operated 393 in full-scan mode during the experiment, and several metabolites were detected and profiled 394 simultaneously as illustrated in Figure 2. Thus, the known metabolites nortriptyline (m/z 264), 395 hydroxy-amitriptyline (m/z 294), and dihydroxy-amitriptyline (m/z 310) were all measured in 396 the experiment. The entire experiment was completed in 45 minutes. 397 9. Technical and operational challenges \u2013 part II 398 In most of our experimental work with c-EME, we have used NPOE as SLM. NPOE is an excellent 399 SLM for non-polar basic drug substances because (a) this solvent provides high extraction 400 recoveries, and (b) because it is highly stable in the porous polypropylene supports. Because of 401 the latter, SLMs based on NPOE can be used for long periods of time without replenishment. In 402 some cases, we have used the same SLM of NPOE for 2-3 days without any interruption (8 hours 403 per day). Unfortunately, NPOE is much less efficient for basic compounds with log P < 2, and for 404 acidic compounds. From experiments with traditional EME, several SLMs have been developed 405 for the aforementioned compounds [10], but many of them are not appropriate for long term 406 use under dynamic conditions due to leakage from the support membrane. Recently, we used a 407 mixture of NPOE and triphenyl phosphate (TPP) (70:30 w/w) for extraction of polar basic 408 drugs, and this SLM showed excellent long term stability [17]. TPP is a solid chemical with very 409 low water solubility, and is therefore not prone to leakage. TPP facilitated transfer of polar basic 410 drugs based on hydrogen bonding interactions. Although TPP can open up for applications for 411 polar compounds, more research is definitely required to increase the number of alternatives. 412 Another challenge we have experienced is illustrated in Figure 5 (data not previously 413 published) using our new highly flexible c-EME flow-probe with make-up flow [17]. When a 414 buffer solution with drug substance (vortioxetine) was pipetted into the sample chamber, the 415 signal measured by the mass spectrometer increased as expected, although there was a small 416 time delay. This delay was approximately 0.30 minutes, and was inherently related to the 417 extraction kinetics and the transfer to the mass spectrometer. The signal increased during 0.6 418 minutes and was then stable because no reaction occurred in the sample chamber. However, 419 when the buffer solution with drug substance was removed momentarily after 10.0 minutes and 420 replaced with pure buffer (no drug substance), the mass spectrometric signal took several 421 minutes to re-establish at background level. After a delay of approximately 0.4 minutes (at 10.4 422 minutes), the mass spectrometric signals decreased rapidly and within 1.0 minute (at 11.4 423 minutes) the signal intensity was reduced from 2x106 to 0.6x16. From this point forward, 424 however, still approximately 4 minutes of operation with pure buffer (at 400 V) was required 425 before the signal re-established at background level. To some extent, this was due to adsorption 426 of drug substance in the reaction chamber and in the connecting tubes, but we have evidence 427 that suggests that drug substances may be partially trapped in the SLM. This, we hypothesize, is 428 due to partial deprotonation of the substances inside the SLM, and may be a general challenge in 429 EME. If some of the molecules are deprotonated, they are no longer influenced by the electrical 430 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 12 field and are slowly released from the SLM only by passive diffusion. From our experiences, the 431 following parameters affect the clearance of drug substances from the SLM: 432 \u2022 Chemical composition of SLM 433 \u2022 Membrane thickness 434 \u2022 pH in sample 435 \u2022 Voltage 436 We have experienced that clearance of drug substance from the SLM is generally improved by 437 reducing the membrane thickness (reduced diffusion distance to acceptor solution), decreasing 438 pH in the sample (increased level of analyte protonation), and by increasing the extraction 439 voltage (improved distribution according to equation 1 and reduced residence time in SLM). 440 These observations are in accordance with the hypothesis of partial deprotonation. The data 441 shown in Figure 5 were obtained with bis(2-ethylhexyl) phosphite (DEHPi) as SLM, the 442 membrane thickness was 150 \u00b5m, sample pH was 3, and voltage was 400 V. The issue around 443 partial deprotonation will be studied in more details in the near future. 444 A final challenge was experienced when moving from micro-chip and into probe systems. While 445 the active acceptor volume was about 600 nL in our micro-chip systems, it was reduced down to 446 90 nL with the probe system. Down-scaling the active acceptor is interesting, and recently 100-447 times enrichment was reported for 1 minute extraction into 8 nL active acceptor [18]. However, 448 experiences gained recently have indicated that down-scaling is challenging, and EME 449 performance may be affected by surface adsorption or related phenomena under nL operation. 450 This is probably due to the increasing surface area-to-acceptor volume ratio inherent with 451 down-scaling. One example of this is illustrated in Table 2 (data not previously published), 452 where the basic drug substances haloperidol and loperamide were extracted under nL 453 operation and subsequently analyzed by capillary electrophoresis. As seen from the data, the 454 extraction of loperamide was enhanced by the presence of other drug substances in the sample, 455 most probably because the latter reduced the surface adsorption of loperamide. Therefore, our 456 current and near-future activities on c-EME will be conducted in the micro-chip format. We 457 expect the micro-chip format to be more robust, and to also show a higher potential for future 458 mass production. 459 10. Related publications 460 The discussions above have been focused on our own work and experiences [11-17]. It should 461 be emphasized that related papers have been published by other research groups with focus on 462 micro-chip configuration [21-23] and on the impact of dynamic conditions (flow of sample and 463 acceptor) [24,25]. The systems were not connected directly to mass spectrometry, but acceptor 464 solutions were collected and analyzed after EME by HPLC or LC-MS. Thus, continuous on-line 465 measurements were not reported, but due to similarity with c-EME, the papers are briefly 466 discussed in the following. In one paper, an EME chip similar to Figure 3 was developed and 467 tested with amitriptyline and nortriptyline as model analytes and urine as sample matrix [21]. 468 The sample volume was 1 mL, and this was pumped into the micro-chip at 30 \u00b5l min-1. The 469 acceptor solution was stagnant and the volume was 20 \u00b5L. After EME the acceptor solution was 470 removed and analyzed by HPLC. Pre-concentration by a factor of 17 was reported, 471 corresponding to a calculated recovery of 30 %. In two subsequent papers, this micro-chip was 472 developed further by implementation of two separate extraction cells in series within the same 473 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 13 embodiment [22,23]. In the first paper, non-polar basic drugs were extracted in one cell, while 474 polar basic drugs were extracted in the second cell [22]. Again, the acceptor solution (total 475 volume of 30 \u00b5L) was stagnant, and subsequently analyzed by HPLC. In the second paper, basic 476 drugs were extracted in the first cell, while acidic drugs were extracted in the second cell. In all 477 three micro-chip papers, the micro-chip channels for EME were of 30 mm length, 200 \u00b5m depth, 478 and 1 mm width. Thus, the active volumes of sample and acceptor solution were approximately 479 6 \u00b5L, and the active surface of the SLM was 30 mm2. Dimensions were slightly larger than in our 480 micro-chip (Figure 3), but performance appeared to be comparable. Similar data reported by 481 two independent research groups is a strong support for EME in the micro-chip configuration. 482 High performance was due to short diffusion distances and due to flow conditions. The 483 advantage of working under flow conditions has been further supported by two alternative 484 approaches to dynamic EME using neuropeptides and antidepressants as model analytes 485 [24,25]. 486 11. Future perspectives 487 The current tutorial has discussed different approaches to continuous electromembrane 488 extraction (c-EME), both from stagnant and dynamic samples. This includes micro-chip, probe, 489 and flow-probe EME. c-EME shows great potential for in-vitro drug metabolism studies, but 490 extraction kinetics and reaction mixture depletion may challenge such experiments. Also, highly 491 stable SLMs have to be developed for polar basic substances and for acidic substances. This is 492 highly important for drug metabolism studies, because many drug metabolites are polar and 493 with acidic functionalities (such as glucuronides). Our research will definitely be focused on c-494 EME in the future, to develop even better technical devices, and to develop new applications. We 495 consider the micro-chip format to be more robust and this show a higher potential for future 496 mass production. Therefore, we will give priority to the micro-chip c-EME. The applications can 497 be pharmaceutical or biomedical, but also other types of reactions in aqueous solution may be 498 monitored using c-EME. c-EME may also be used in combination with different type of sensors, 499 where continuous sample clean-up is required in order to eliminate possible interferences from 500 the sample matrix. 501 502 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 14 References 503 [1] S. Pedersen-Bjergaard, K.E. Rasmussen, Electrokinetic migration across artificial liquid membranes: new 504 concept for rapid sample preparation of biological fluids, J. Chromatogr. A 1109 (2006) 183-190. 505 [2] P. Kuban, A. Slampova, P. Bocek, Electric field-enhanced transport across phase boundaries and membranes 506 and its potential use in sample pretreatment for bioanalysis, Electrophoresis 31 (2010) 768-785. 507 [3] V.K. Marothu, M. Gorrepati, R. Vusa, Electromembrane Extraction\u2014A Novel Extraction Technique for 508 Pharmaceutical, Chemical, Clinical and Environmental Analysis, J. Chromatogr. Sci. 51 (2013) 619-631. 509 [4] A. Gjelstad, S. Pedersen-Bjergaard, Recent developments in electromembrane extraction, Anal. Method. 5 510 (2013) 4549-4557. 511 [5] P.W Lindenburg, R. Ramautar, Thomas Hankemeier, The potential of electrophoretic sample pretreatment 512 techniques and new instrumentation for bioanalysis, with a focus on peptidomics and metabolomics, 513 Bioanalysis 5 (2013) 2785-2801. 514 [6] A. Gjelstad, S. Pedersen-Bjergaard, Electromembrane extraction \u2013 Three-phase electrophoresis for future 515 preparative applications, Electrophoresis 35 (2014) 2421-2428. 516 [7] C. Huang, K.F. Seip, A. Gjelstad, S. Pedersen-Bjergaard, Electromembrane extraction for pharmaceutical and 517 biomedical analysis \u2013 Quo vadis, J. Pharm. Biomed. Anal. 113 (2015) 97-107. 518 [8] M. Rezazadeh, Y. Yamini, S. Seidi, Electrically stimulated liquid-based extraction techniques in Bioanalysis, 519 Bioanalysis 8 (2016) 815-828. 520 [9] A. Oedit, R. Ramautar, T. Hankemeier, P.W. Lindenburg, Electroextraction and electromembrane extraction : 521 Advances in hyphenation to analytical techniques, Electrophoresis 37 (2016) 1170-1186. 522 [10] C. Huang, A. Gjelstad, S. Pedersen-Bjergaard, Organic solvents in electromembrane extraction: recent 523 insights, Rev. Anal. Chem. 35 (2016) 169-183. 524 [11] N.J. Petersen, H. Jensen, S. Honor\u00e9 Hansen, S. Taule Foss, D. Snakenborg, S. Pedersen-Bjergaard, On-chip 525 electro membrane extraction, Microfluid. Nanofluid. 9 (2010) 881-888. 526 [12] N.J Petersen, S. Taule Foss, H. Jensen, S. Honor\u00e9 Hansen, C. Skonberg, D. Snakenborg, J.P. Kutter, and S. 527 Pedersen-Bjergaard, On-chip electro membrane extraction with on-line ultraviolet and mass spectrometric 528 detection, Anal. Chem. 83 (2011) 44-51 529 [13] N.J Petersen, J. S\u00f8nderby Pedersen, N. N\u00f8rg\u00e5rd Poulsen, H. Jensen, C. Skonberg, S. Honor\u00e9 Hansen, and S. 530 Pedersen-Bjergaard, On-chip electromembrane extraction for monitoring drug metabolism in real time by 531 electrospray ionization mass spectrometry, Analyst 137 (2012) 3321-3327. 532 [14] H. Bonkerud Dugstad, N.J. Petersen, H. Jensen, C. Gabel-Jensen, S. Honor\u00e9 Hansen, and S. Pedersen-Bjergaard, 533 Development and characterization of a small electromembrane extraction probe coupled with mass 534 spectrometry for real-time and online monitoring of in vitro drug metabolism, Anal. Bioanal. Chem. 406 535 (2014) 421-429. 536 [15] D. Fuchs, H. Jensen, S. Pedersen-Bjergaard, C. Gabel-Jensen, S. Honor\u00e9 Hansen, N.J Petersen, Real time 537 extraction kinetics of electro membrane extraction verified by comparing drug metabolism profiles 538 obtained from a flow-flow electro membrane extraction-mass spectrometry system with LC-MS. Anal. Chem. 539 87 (2015) 5775-5781. 540 [16] D. Fuchs, C. Gabel-Jensen, H. Jensen, K.D. Rand, S. Pedersen-Bjergaard, S. Honor\u00e9 Hansen, N.J. Petersen, 541 Direct coupling of a flow-flow electromembrane extraction probe to LC-MS, Anal. Chim. Acta 905 (2016) 93-542 99. 543 [17] T. Kige Rye, D. Fuchs, S. Pedersen-Bjergaard, N.J Petersen, Direct Coupling of Electromembrane Extraction to 544 Mass Spectrometry - Advancing the probe functionality toward measurements of zwitterionic drug 545 metabolites, Anal. Chim. Acta 983 (2017) 121-129 546 [18] M.D. Ramos Pay\u00e1n, B. Li, N.J. Petersen, H. Jensen, S. Honor\u00e9 Hansen, S. Pedersen-Bjergaard, Nano-547 electromembrane extraction, Anal. Chim. Acta 785 (2013) 60-66. 548 [19] K.F. Seip, H. Jensen, M. Hovde S\u00f8nsteby, A. Gjelstad, S. Pedersen-Bjergaard, Electromembrane extraction: 549 Distribution or electrophoresis?, Electrophoresis 34 (2013) 792-799. 550 [20] A. Gjelstad, K.E. Rasmussen, S. Pedersen-Bjergaard, Simulation of flux during electromembrane extraction 551 based on the Nernst-Planck equation, Anal. Chim. Acta 1174 (2007) 104-111. 552 [21] Y. Abdossalami Asl, Y. Yamini, S. Seidi, B. Ebrahimpour, A new effective on chip electromembrane extraction 553 coupled with high performance liquid chromatography for enhancement of extraction efficiency, Anal. Chim. 554 Acta 898 (2015) 42-49. 555 [22] Y. Abdossalami Asl, Y. Yamini, S. Seidi, A novel approach to the consecutive extraction of drugs with different 556 properties via on chip electromembrane extraction, Analyst 141 (2016) 311-318. 557 [23] Y. Abdossalami Asl, Y. Yamini, S. Seidi, M. Rezazadeh, Simultaneous extraction of acidic and basic drugs via 558 on-chip electromembrane extraction, Anal. Chim. Acta 937 (2016) 61-68. 559 [24] N. Drouin, S. Rudaz, J. Schappler, Dynamic-electromembrane extraction : A technical development for the 560 extraction of neuropeptides, Anal. Chem. 88 (2016) 5308-5315. 561 [25] Y.A. Asl, Y. Yamini, S. Seidi, H. Amanzadeh, Dynamic electromembrane extraction: Automated movement of 562 donor and acceptor phases to improve extraction efficiency, J. Chromatogr. A 1419 (2015) 10-18. 563 564 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 15 Figure captions 565 566 Figure 1 Schematic illustration of a) probe, b) micro-chip, and c) flow-probe c-EME 567 568 Figure 2 In-vitro metabolism of amitriptyline measured by micro-chip c-EME and mass 569 spectrometry. Amitriptyline was monitored at m/z 278. NT, AT-OH, and AT-2(OH) are 570 nortriptyline (m/z 264), hydroxy-amitriptyline (m/z 294), and dihydroxy-amitriptyline (m/z 571 310), respectively. Additions to the reaction mixture were a) amitriptyline, b) rat liver 572 microsomes, and c) co-factor [13] 573 574 Figure 3 Schematic illustration and photo of micro-chip c-EME coupled directly to electrospray 575 ionization mass spectrometry for in-vitro drug metabolism studies [13] 576 577 Figure 4 Schematic illustrations of a) probe and b) flow-probe c-EME coupled directly to 578 electrospray ionization mass spectrometry [14,15] 579 580 Figure 5 Flow-probe c-EME and mass spectrometry from buffer solution with vortioxetine (m/z 581 299). The buffer solution with vortioxetine was removed and replaced by pure buffer solution 582 after 10.0 minutes (data not previously published) 583 584 585 586 ACCEPTED MANUSCRIPT 16 590 Figure 1 591 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 17 592 Figure 1, continued 593 594 595 596 597 598 599 600 601 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 18 602 603 Figure 2 604 605 606 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 19 607 608 609 Figure 3 610 611 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 20 612 613 614 615 Figure 4 616 617 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT 21 618 Figure 5 619 620 0 500000 1000000 1500000 2000000 2500000 3000000 3500000 0 5 10 15 20 400 V M ANUSCRIP T ACCEPTE D 22 ______________________________________________________________________________________________________________ 622 Sample Metabolic reaction mixture 623 Sample flow-rate 20 \u00b5L min-1 624 Acceptor solution 10 mM HCOOH 625 Acceptor solution flow-rate 20 \u00b5L min-1 626 Supported liquid membrane 2-nitrophenyl octyl ether (NPOE) 627 Voltage 15 V 628 ______________________________________________________________________________________________________________ 629 630 Table 2 Nano-EME of haloperidol and loperamide in presence of other drug substances1 631 ______________________________________________________________________________________________________________ 632 Peak area (\u00b1 standard deviation, n=4) 633 _________________________________________________________ 634 Haloperidol Loperamide 635 _____________________________________________________________________________________________________________ 636 637 No other drug substances in the sample 13,644 \u00b1 1,682 40,844 \u00b1 7,325 638 639 Pethidine, nortriptyline, and methadone 14,374 \u00b1 632 76,697 \u00b1 4,671 640 present in sample 641 _____________________________________________________________________________________________________________ 642 1All drug concentrations were 1 ppm 643 644 M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT Highlights \u2022 Continuous electromembrane extractions is discussed \u2022 Principles, theory, technical development, operational parameters, performance, applications, and future perspectives are highlighted \u2022 Discussion based on both published and unpublished work M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT Cristina Rom\u00e1n Hidalgo is a PhD student from the University of Seville (Spain). She develops her doctorate through a \u201cV Plan Propio de Investigaci\u00f3n\u201d fellowship of the University of Seville. She graduated in Chemistry in 2013 and got a \u201cMaster in advanced studies in Chemistry\u201d in 2014. She has two publications in peer-reviewed international journals. M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT David Fuchs is Post-Doctoral researcher at the Department of Medical Biochemistry and Biophysics at the Karolinska Institute, Stockholm (Sweden). He obtained his PhD at the Department of Pharmacy, University of Copenhagen (Denmark) where he focused his research on elucidating the potential of electromembrane extraction for automated and high-throughput sample preparation applications. M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT Henrik Jensen is Associate Professor at the Department of Pharmacy, University of Copenhagen (Denmark). He research focus is on analytical chemistry and pharmaceutical physical chemistry and on developing numerical and mathematical model for in depth understanding of new technologies. He has published more than 110 peer reviewed papers, patents and book chapters. M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT J\u00f6rg P. Kutter has more than 20 years of research and teaching experience with microchip-based analytical tools. Since September 2013, he is the Chair of Analytical Biosciences at the Dept. of Pharmacy at the University of Copenhagen. His research interests focus on the development of microfluidic devices for applications in the life sciences, and in particular the pharmaceutical sciences. He is published over 100 scientific articles. M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT Maria Ramos Pay\u00e1n is Senior Researcher at Microelectronic National Centre of Barcelona and Associate Professor at Univertitat Autonoma of Barcelona, Spain. She is specialized in microextraction procedures for sample treatment and microfluidic devices applied in environmental, biological and biomedical field. She has published more than 25 papers in international journals. M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT Nickolaj Jacob Petersen is Associate Professor at the Department of Pharmacy, University of Copenhagen (Denmark). His expertise comprises miniaturized systems (\u00b5-TAS) and electrokinetic transport for sample preparation as well as on electrokinetic separation methods coupled to mass spectrometry. He has published more than 25 peer reviewed papers. M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT Stig Pedersen-Bjergaard is Professor at School of Pharmacy, University of Oslo (Norway), and Professor (part time) at Department of Pharmacy, University of Copenhagen (Denmark). He has specialized in analytical micro extraction technologies, on development and applications of artificial liquid membranes, and on electrokinetic separation methods. He has published more than 170 papers in international journals. M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT M ANUSCRIP T ACCEPTE D ACCEPTED MANUSCRIPT"
        ]
    },
    {
        "image_filename": "designv11_13_0001997_j.phpro.2014.08.081-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001997_j.phpro.2014.08.081-Figure1-1.png",
        "caption": "Fig. 1. (a) Simplified scheme of welding process boundaries; (b) clamping fixture.",
        "texts": [
            " It is known that at higher feeding rates melt pool instabilities such as grooving and/or humping occur. In order to determine the laser micro welding process, ultra-thin ( 100 \u03bcm) stainless steel (AISI 304) foils served as specimens. Applications in micro welding are frequently jointed in a lap weld (Qin, 2010). For this reason, the foils were welded in an overlapping joint. The feeding rate was raised at a given laser power until an incomplete penetration resulted. Thereby, a process window could be determined (Fig. 1(a)). This step was iterated by increasing the laser power incrementally. If only results of the lower process boundary nearby an incomplete penetration were used, this method of approach leads to the following definition: The welding depth s is in accordance with the foil thickness and then defined by it. Two different laser sources were used in this study in order to provide a broad range of focal diameters. The first laser source is a ROFIN multi-mode fiber laser which was coupled to a RAYLASE scanning unit",
            " In combination with different beam properties of the laser sources, focal diameters between 25 \u03bcm and 204 have been obtained. * Measured at 10 % of maximum power level A beam profile measurement with all optical settings was carried out, and focal diameters as well as beam quality factors were determined. The applied power on the work piece was measured. In order to avoid a focal shift due to thermal lensing, the processing time was limited. Welding paths with a length of approximately 50 mm were applied on the clamped foils. The free clamping length was 8 mm (see Fig. 1(b)). 3. Process windows The upper and lower process boundaries of two different welding depths (50 \u03bcm and 100 \u03bcm) were determined and displayed in the following figures. First of all, in Fig. 2(b) can be seen that the lower process boundary of focal diameters of 25 \u03bcm and 31 \u03bcm and a welding depth of 50 \u03bcm proceeds discontinuously. At a power level above 300 W the maximum achievable feeding rate decreases due to a surplus supply of power. Splatter formation increases and the melt is driven out (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000182_msf.765.413-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000182_msf.765.413-Figure1-1.png",
        "caption": "Fig. 1. Schematic representation of the reference axes.",
        "texts": [
            " A MTT SLM 250 equipment was used to produce the SLM samples, with a laser power of 175 W, a travel speed of 710 mm/s and with Ar flowing systematically in the Ox direction. The lasing strategy was complex, imposing a rotation of 79\u00b0 between two successive layers. In the EBM process, on the other hand, the samples have been made using a random scanning strategy [11]. The titanium substrate was preheated at the temperature of 200 \u00b0C for SLM and 650 \u00b0C for EBM, whereas the substrate had not been pre-heated in the case of the LC process. Elongated samples for SLM and EBM have been oriented along the three reference axis as presented in Fig. 1a while samples have all been made along the Oz axis in the LC process, as illustrated in Fig. 1b. In all techniques, the building direction is oriented vertically along the Oz axis. The samples metallurgical health has been assessed as sound (no crack and a volume fraction of porosity under 0.5 %). The samples have been polished (following standard practices) and etched with Kroll\u2019s reagent (i.e. 5 % HNO and 5 % HF in distilled water) to reveal their microstructures which were then observed using an Olympus BX60M optical microscope. Laser Cladding. Fig. 2 shows three different sections of a sample made by LC",
            " From these, a direct correlation can be drawn between the intensity of the thermal gradient and the grain size: the greater the thermal gradient, the smaller the width of the primary grains. Fig. 3a shows the structure formed upon cooling the primary grains, exhibiting the morphology typical of a martensitic microstructure. Indeed, in the case of LC process, the temperature drops so quickly that the primary would undergo a displacive transformation [8]. Selective Laser Melting. Fig. 4 presents the cross sections of SLM samples processed according to the orientation described in Fig. 1a. Fig. 4a) shows the plane perpendicular to the building direction which perpendicularly intersects the elongated grains apparent in Fig. 4b. Fig. 4a exhibits a rather equiaxed morphology with a grain diameter of approximately 50 \u00b5m. It is, moreover, worth noting that the grain width does not change significantly along the sample height as observed previously in LC samples (Figs. 2b&c). This absence of evolution of the grain width along the height could be explained by a decrease of the thermal gradient due to the preheating of the substrate at a temperature of 200 \u00b0C",
            " But these elongated grains are not apparent on Fig. 4c. This suggests that, for some samples orientations, the primary \u03b2 grains are tilted with respect to the building direction (i.e. the oz axis). It has been suggested [10,12] that this could be due to a combined effect of part geometry and of evaporation phenomena caused by the continuous Ar flow in the Ox direction. Electron Beam Melting. Fig. 5 presents the cross sections of the three EBM samples processed with different orientation as illustrated in Fig. 1a. Fig. 5a shows the plane perpendicular to the building direction, thus intersecting perpendicularly with the elongated primary \u03b2 grains as previously noted for the SLM samples (Fig. 4a). Fig. 5b&c further shows that the primary \u03b2 grains are elongated following the building direction, suggesting that, in the case of EBM contrarily to SLM, these grains were not tilted, which could be ascribed to the more random scanning strategy [11]. Moreover, preheating of the substrate up to 650 \u00b0C leads to a greater thermal homogeneity throughout the part"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002485_s00170-016-8430-x-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002485_s00170-016-8430-x-Figure9-1.png",
        "caption": "Fig. 9 a Case 2: ECE angles taken equal to the respective flank angles. b Tilt corrections",
        "texts": [
            " Due to this condition, the sum of ECE angles is equal to the sum of gear flank angles, and the two conditions are fulfilled. The two coincident profiles of the cutter sweep surface and the gear tooth are shown in the Fig. 8. In industry, standard blade angles are used, therefore, standard blade angles or equal ECE angles are considered in this case, such that the sum of ECE angles should be equal to the sum of gear flank angles, hence the symmetric axes s!e and s!g coincide with each other as shown in the Fig. 9a. In this case, the coordinate systems Ob of the cutter sweep surface and the pointN of gear teeth profile are coincided, such that the cutter is tilted about the Yb axis with a tilt angle of \u03b8g as shown in the Fig. 9a. Angle \u03b8g is measured from the s! axis of the gear tooth profile to the Xb axis of the ECE. Mathematically, it is given by \u03b8g \u00bc \u0394\u03b1g\u2212\u0394\u03b1e: \u00f016\u00de Due to the tilt of cutter-head, the tangency of the two surfaces at the mean point projections does not exist and cutter gauges inside the root cone of the gear, either at the convex or concave side of the gear teeth flank. To eliminate this problem, cutter location corrections are calculated, and the cutter is displaced along the s! axis and point width of the gear with the corrections xc and zc, respectively. These corrections provide the tangency of cutter sweep and tooth surfaces at least at convex side of the gear tooth flank, which is necessary due to the higher compressive forces between the gear and the pinion [2]. The corrected cutter engagement is show in the Fig. 9b. Due to the corrections, the symmetric axes s!e and s!g does not coincide but remain parallel to each other, but the first condition is fulfilled. For positive \u03b8g angle, the correction xc is along the upward direction, whereas zc is along the righthand side, whereas direction of only zc reverse, in case of negative \u03b8g angle. These corrections can be calculated by xc \u00bc 0:5\u22c5Pw\u22c5sin \u03b8g ; \u00f017\u00de zc \u00bc 0:5\u22c5Pw\u22c5 1\u2212cos\u03b8g \u22c5 \u03b8g . \u03b8g : \u00f018\u00de In Section 3, relationship between the geometry of the gear teeth and the cutter sweep surface is built, whereas in this section, an algorithm for the determination of the inside and outside blade angles\u03b1b,k is presented"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000787_j.ifacol.2015.10.272-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000787_j.ifacol.2015.10.272-Figure2-1.png",
        "caption": "Fig. 2. ROV propellers system.",
        "texts": [
            " The quantities Tx, Ty and Mz appearing in (17), (18) and (19) are the decomposition of the thrust and the torque provided by the four vehicle propellers along the axes ofR; the corresponding decomposition with respect to the axes of Ra will be denoted with Txa, Tya and Mza. For the considered case with ROV operating on surfaces parallel to the x \u2212 y plane, the two vectors \u03c4 = [Tx Ty Mz] T and \u03c4 a = [Txa Tya Mza] are related by the following relationship \u03c4 = \ufffd c\u03c8 \u2212s\u03c8 0 s\u03c8 c\u03c8 0 0 0 1 \ufffd \u03c4 a (20) and, for the considered ROV with four propellers disposed as shown in Figure 2, ones has \u03c4 a = \ufffd c\u03b1 c\u03b1 c\u03b1 c\u03b1 \u2212s\u03b1 \u2212s\u03b1 s\u03b1 s\u03b1 \u2212da da \u2212da da \ufffd u (21) with u = [T1 T2 T3 T4] T , \u03b1 = \u03c0/4 and da = (dxs\u03b1 + dyc\u03b1). Thus, for the considered ROV, where the control inputs are more than the degrees of freedom, the thruster configuration matrix, reported in (10), is not square, and can be expressed as B = \u23a1 \u23a3 c(\u03c8\u2212\u03b1) c(\u03c8\u2212\u03b1) c(\u03c8+\u03b1) c(\u03c8+\u03b1) s(\u03c8\u2212\u03b1) s(\u03c8\u2212\u03b1) s(\u03c8+\u03b1) s(\u03c8+\u03b1) \u2212da da \u2212da da \u23a4 \u23a6 (22) August 24-26, 2015. Copenhagen, Denmark Antonio Fasano et al. / IFAC-PapersOnLine 48-16 (2015) 146\u2013151 149 The model (10) does not take into account possible failures or faults of the thrusters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000999_j.proeng.2017.02.307-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000999_j.proeng.2017.02.307-Figure4-1.png",
        "caption": "Fig 4. (\u0430) test bench without a vibroacoustic cover: 1- pump, 2- motor, 3- frequency converter, 4- inlet and outlet pipes; (b) test bench with a vibroacoustic cover: 5- manometer, 6- vibroacoustic cover.",
        "texts": [
            " Gears from different polymer composite materials have been manufactured by machining (Figure 3). Fig 2. The gear micropump case from PEEK (a) and stainless steel (b). Fig 3. Driver shaft from different materials. 3. Experiments The research has been conducted on an existing test bench [24]. The test bench allows setting up and mixing several parameters such as outlet pressure, speed of shaft rotation, flow rate, torque, and fluid temperature. The hydraulic oil HLP-46 has been used. The overview of the test bench is presented in Figure 4. Vibroacoustic data acquisition was performed with special equipment [25]. The noise-level meter \u201cEcophizika110A\u201d designed to collect vibroacoustic data at octave, 1/3 octave, and narrow band range (Fig. 5). A threecomponent accelerometer (Fig. 5) provides an opportunity to collect all three vibration components. The \u201cEcophizika-110A\u201d technical features are presented in table 1. Several rotors packages made of different polymer composite materials were studied on a gear micropump. A number of modes were formed as a combination of outlet pressure and rotational shaft speed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001009_iet-epa.2016.0680-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001009_iet-epa.2016.0680-Figure3-1.png",
        "caption": "Fig. 3 Field-orientated reference frame of IM and IFOC controller",
        "texts": [
            " Then, as can be seen from (14) and (15), if L\u2217m is not equal to Lm, the rotor flux and torque will deviate from actual value of IM. Another reason, actually resulted in mutual inductance mismatch, is rotor flux cr orientation deviation. Although the synchronous speed in controller w\u2217 e and IM we are equivalent in steady-state, they are unequal in transient state, causing their integration h\u2217 r and hr unequal. That means the \u2018rotor flux orientation d\u2217-axis\u2019 in controller does not overlap with the actual cr orientation d-axis. The angle deviation is denoted by u, as shown in Fig. 3. u can be both positive and negative, related to the sign of (L\u2217m \u2212 Lm), which will be proved in the next section. The d\u2217 \u2212 q\u2217 represent coordinate axes of controller (black solid lines), while d \u2212 q represent axes of motor (grey dashed lines). Reference current components i\u2217sd and i\u2217sq are calculated from (14) and (15). As shown in Fig. 3, i\u2217sd in d\u2217 axis can decompose into isdq and isdd in d \u2212 q coordinate system. Similarly, i\u2217sq can decompose into isqd and isqq in d \u2212 q frame. When u . 0, as illustrated in Fig. 3, the component isqd , decomposed from the expected torque-producing current i\u2217sq, is actually working as magnetising current component. In the same time, the component isdq, from expected magnetising current i\u2217sd , is actually in torque-producing direction. Furthermore, its direction is opposite to isqq, which causing the total torque-producing current isq even smaller than i\u2217sq. This is failure of decoupling of magnetising and torque-producing currents, which will cause torque and flux steady-state errors",
            " One is directly influences the reference currents calculation by (14) and (15). Another one is influences errors indirectly by causing orientation deviation, leading to failure of currents decoupling. Steady-state errors in this paper include orientation error u, rotor flux error Dcr and torque error DTe . All the calculation formulas will be deduced in the following sections. In IFOC controller, the torque and rotor flux are calculated by T\u2217 e = 3 2 np L\u22172m L\u2217r i\u2217sdi \u2217 sq (20) c\u2217 r = L\u2217mi \u2217 sd (21) Referring to Fig. 3, the current components in d \u2212 q axes of motor 1107 are isd = isdd + isqd isq = isqq \u2212 isdq (22) where isdd = i\u2217sd cos u isdq = i\u2217sd sin u (23) isqd = i\u2217sq sin u isqq = i\u2217sq cos u (24) Since d\u2217 \u2212 q\u2217 and d \u2212 q axes have equal rotation speed w\u2217 e = we, they share the same slip speed w\u2217 sl = wsl in steady state. According to (12), the following equation holds Rr L\u2217r i\u2217sq i\u2217sd = Rr Lr isq isd (25) Combining (22)\u2013(25), the angle u can be deduced as tan u = (i\u2217sq/i \u2217 sd) 1\u2212 (Lr/L \u2217 r ) ( ) 1+ i\u2217sq/i\u2217sd ( )2 (Lr/L \u2217 r ) (26) From the above equation, the angle u is related to reference currents i\u2217sd and i \u2217 sq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000212_detc2013-12809-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000212_detc2013-12809-Figure3-1.png",
        "caption": "FIGURE 3. THE HEAT BAND ON THE DISC FOR A RANDOMLY SELECTED SOLUTION FROM THE PARETO FRONT.",
        "texts": [
            " In order to avoid generating 200 different CAD models manually, one for each thickness, 10 different thickness from 5[mm] to 14[mm] (5,6...14) are selected which for each thickness 20 sampling points are created by using the Latin hypercube function (LHSDESIGN) in MATLAB. Three separate work stations were simultaneously used to run the finite element simulations, one thickness at a time and the total 20 DoE of each thickness consumed approximately 120,000 seconds of computation time. Therefore, all 200 simulations was completed at around 5 days. Figure 3 illustrates the heat band on the surface of the disc brake generated from the finite element model, for one random solution selected from the Pareto front. The maximum temperature in terms of number of increments used in the finite element simulation, for the the same randomly selected point is plotted in Fig. 4. All of the total 200 sampling points and their related objective function values are used to construct two surrogate models for the first two objectives, by employing the RBFN with a priori bias method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001136_s00170-017-0625-2-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001136_s00170-017-0625-2-Figure4-1.png",
        "caption": "Fig. 4 Scanning strategy of the ladder block structure: a scanning strategy and b sketch of overlap and deposition path. a The deposition directions are indicated as arrows. b The deposition path of the beads in a particular layer is indicated by a red arrow on the right side",
        "texts": [
            " Eight different thin-wall samples were deposited with 1, 2, 3, 5, 8, 13, 21, and 34 layers, which are referred to as S1, S2, S3, S5, S8, S13, S21, and S34, respectively. The processing parameters for the thin wall sample are recorded in Table 2. Moving distance of the Z-axis in each layer in S1\u2013S34 samples is recorded in Table 3. In this study, the shapes of molten pools vary with the deposition process. Therefore, the increase in height of depositions are different in each layer of depositions, which will be lower in the first few layers and be stable after nine layers of the depositions, as shown in Table 3. Figure 4 shows the scanning strategy of the ladder block structure. The processing parameters for the ladder block were the same as those for the thin wall, as listed in Table 2. Six different ladder block samples were deposited with 1, 2, 3, 5, 8, and 13 layers, which are represented as X1, X2, X3, X5, X8, and X13, respectively. The scanning path is always along the same direction from the beginning of each layer, as depicted in Fig. 4a, in which the deposition directions are indicated as arrows. The overlap ratio for each adjacent bead is 0.5, and the beads in a particular layer are always deposited exactly over the corresponding beads of the previous layer, as shown in Fig. 4b. Herein, the \u201coverlap ratio\u201d is recorded as OVr, and calculated using Eq. 1: OVr \u00bc OL=W \u00f01\u00de where OL is overlap width and W is single bead width, as shown in Fig. 4b. Furthermore, for recording the temperature history of the molten pool in the deposition process, infrared thermography was applied. The corresponding method is illustrated in Fig. 5. The FLIRX6530sc was the infrared thermo-graph used in this study. Figure 6 shows the different sections for metallographic observation, which are represented as sections X and Y. Section X is normal to the scanning direction, while section Y is parallel to the scanning direction. All the cross-sections were made by wire cutting"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000335_j.euromechsol.2012.10.011-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000335_j.euromechsol.2012.10.011-Figure1-1.png",
        "caption": "Fig. 1. (a) A guidewire passes through the lesion area of an artery channel. It is assumed that the guidewire is squeezed by the elastic tissue in the lesion. (b) The tissue in the lesion is modeled as a pair of pads suspended elastically by translational springs a and torsional springs b. The translational springs are pre-compressed by force F1.",
        "texts": [
            " Stent procedures are commonly adopted in treating patients with coronary artery diseases (Schneider, 2003). During the surgery, the surgeon inserts a guidewire into the artery of the patient through a puncture in the groin. With the help of a fluoroscopic imaging system, the guidewire is pushed to the neighborhood where artery repair is needed. In some extreme cases, the blockage of the artery is so severe that the leading end of the guidewire can only barely pass through the lesion area, as shown in Fig. 1(a). In this case the narrowed passageway may \u201csqueeze\u201d the guidewire and exert a friction force opposing its movement. The guidewire may then deform and contact the artery wall while it is continuously fed into the artery channel with the input end controlled by the surgeon\u2019s hand movement. As a consequence, the movement of the leading end of the guidewire near the lesion does not follow the surgeon\u2019s handmovement at the input end. This may create a problem for a surgeon in the operating room, and for a control engineer in designing a medical robot. This paper intends to investigate the deformation of the guidewire in this situation, son SAS. All rights reserved. with emphasis on the relative movement between the leading end and the input end. The guidewire is modeled as an elastica and the artery housing is modeled as a rigid channel. The lesion area is modeled as a pair of clamps suspended by torsional and translational springs, as shown in Fig. 1(b). The elastica problem discussed in this paper falls in a field called constrained elastica, which refers to an elastica constrained laterally and forced to deform between a pair of rigid walls. Most of the previous researches in constrained elastica assume that the contact between the elastica and the rigid walls is frictionless; see Feodosyev (1977), Vaillette and Adams (1983), Adams and Benson (1986), Adan et al. (1994), Domokos et al. (1997), Holmes et al. (1999), Chai (1998), Chen and Li (2007), and Ro et al",
            " The elastic beam is fed in the channel through the hole without friction and clearance. At the right end B of the channel, there is an elastically suspended clamp. The distance between ends A and B is L. The elastic beam is guided to pass through between the upper and lower rigid pads of the clamp. The pair of pads is supported in the transverse direction by two translational springs a, and rotationally by torsional springs b. The clamp is prevented against moving in the longitudinal direction, as shown in Fig. 1(b). This clamp-spring assembly is meant to represent the stiffness of the blockage tissue inside the artery. In order to simulate the phenomenon that the guidewire be squeezed by the blockage, we assume that the translational springs are precompressed by force F1. The left feeding hole and the right clamp before deformation are at the centerline of the channel. The elastic beam is under edge thrust FA at the left end A. When the squeezing force F1 at the output end B is present, the pair of clamp pads exerts a Coulomb friction force against the longitudinal movement of the beam relative to the clamp"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000199_s00170-014-6551-7-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000199_s00170-014-6551-7-Figure2-1.png",
        "caption": "Fig. 2 Reference systems used in the LT modelling",
        "texts": [
            " The first two parameters correspond to the lengths of the links between successive reference systems, and the other two are rotation angles of one system regarding the other. The D-H model establishes four successive transformations to relate the i frame to the i-1 frame. These four transformations give the homogenous matrix presented by Eq. 4. i\u22121A \u00bc Tz;d \u22c5Rz;\u03b8\u22c5Tx;a\u22c5Rx;\u03b1 \u00bc cos\u03b8i \u2212cos\u03b1i\u22c5sin\u03b8i sin\u03b1i\u22c5sin\u03b8i ai\u22c5cos\u03b8i sin\u03b8i cos\u03b1i\u22c5cos\u03b8i \u2212sin\u03b1i\u22c5cos\u03b8i ai\u22c5sin\u03b8i 0 sin\u03b1i cos\u03b1i di 0 0 0 1 2 664 3 775\u00f04\u00de The LT model developed obtains the reflector location (reference system 3) with respect to the origin of the LT (reference system 0) in terms of \u03b8,\u03c6 and d (see Eq. 5). Figure 2 shows the reference systems used, and Table 2 gives the initial D-H parameters used. 0T 3 \u00bc 0A1\u22c51A2\u22c52A3 \u00f05\u00de The definition of the kinematic model will determine the nature and number of necessary parameters and will allow the accuracy of the system to be improved. The kinematic errors have been modelled in this paper, and the optical errors have been considered in the last kinematic chain as searching for collecting their influence through the kinematic parameters of the last link in the chain rather than explicitly model them"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002720_978-3-319-24055-8-Figure3.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002720_978-3-319-24055-8-Figure3.1-1.png",
        "caption": "Fig. 3.1 The force shapes that are derived from the electro-magnetic simulation are expressed as a spatial Fourier series",
        "texts": [
            " From this the resulting vibration due to a number of excitations and their harmonics can be calculated, and the operating deflected shape can be viewed, to give the resulting vibration at any frequency. The key targets of this methodology are that (i) the speed of the process should be such that the results are available within a time frame that they are useful for making design decisions and (ii) the results can be interpreted easily so as to guide engineering design decisions [16]. With regard to the electric motor, the excitation comes in the form of force shapes that arise from the electromagnetic simulation. Details are given below (Fig. 3.1). 32 B. James et al. The main objective of the electromagnetic simulation is to provide the force-shape amplitudes for the modal superposition. Thus the model of the electrical machine has to calculate the relevant forces in the machine and spatially decompose them into the force shapes. These force shapes are the modal excitation of the structure\u2019s eigenmodes. The forces which influence the acoustic behaviour of the machine are usually radial force. The introduced method, however, works also for tangential and axial forces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure7-1.png",
        "caption": "Fig. 7. Vapor mass fraction changes on the surface of the cylindrical body, under the water\u2013vapor flow.",
        "texts": [
            " The highest force is at the front and at the end part of the body. Total drag force, which influences the motion of the body, is 0,015 N. Due to the vortex beyond the cylindrical body, the drag force at the end part becomes negative. Fig. 5 shows water velocity change nearby the hot body. One can see in Figs. 3 and 5 s a significant difference on the front zone where the flow abruption appears. Fig. 6 shows the change of flow density. Sharp decrease in density appears nearby the cylindrical surface. Fig. 7 shows a change of vapor mass fraction on the surface of the body. Maximum vapor generation is noticed on the front part of the cylinder immediately behind the flow abruption point. Average vapor mass fraction on the surface is 0.008. Further more water is contacted with the body\u2019s surface and cooling becomes more intensive, therefore vapor is destroyed by colder water. At the end part of the cylinder, where the vortex starts and velocity reduces, cooling goes down, and the amount of vapor increases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000458_j.ast.2015.02.007-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000458_j.ast.2015.02.007-Figure2-1.png",
        "caption": "Fig. 2. Rolling and non-rolling frames relation.",
        "texts": [
            " Rotor torque is proportional to winding\u2019s current as \u03c4 = Kt i (9) Exploiting a speed ratio reducer, the relation between output roll angle of rotor and canards deflection has the form: \u03d5 = N\u03b4c (10) Applying PD controller, vm(t) = K P (\u03b4r \u2212 \u03b4c) + K D d(\u03b4r \u2212 \u03b4c) dt (11) and assuming free flight phase, \u03b4r = 0, neglecting the small inductance of the circuit and combining Eqs. (6)\u2013(11) yield, G(s) = \u03b4c(s) \u03b1(s) = Ks T 2 s s2 + 2\u03bcs Tss + 1 (12) where Kt K P N \u2212 RCh (13) Ts = \u221a R( J R N2 + J L) Kt K P N \u2212 RCh (14) \u03bcs = Kt N(Kb N + K D) 2 \u221a R( J R N2 + J L)(Kt K P N \u2212 RCh) (15) The actuator\u2019s block diagram is shown in Fig. 1. In the body-fixed frame, transfer function between actuator de- flection and angle of attack and side slip are:\u23a7\u23aa\u23aa\u23a8\u23aa\u23aa\u23a9 \u03b4y(s) \u03b1(s) = G(s) \u03b4z(s) \u03b2(s) = \u2212G(s) (16) Considering \u03be = \u03b2 + j\u03b1, and \u03b4 = \u03b4y + j\u03b4z , thus \u03b4(s) = \u2212 j\u03be(s)G(s) (17) As Fig. 2 demonstrates, rolling and non-rolling frames are related by each other through [18] \u03b4 = \u03b4\u0303e\u2212 j\u03c6, \u03be = \u03be\u0303e\u2212 j\u03c6 (18) In the non-rolling frame, \u03b4r and \u03b1 as reference deflection and angle of attack are inputs of block diagram and the canard deflection \u03b4c is output. Substituting Eq. (12) into Eq. (17), the time domain of this equation can be described as T 2 s \u00a8\u0303 \u03b4 + ( 2\u03bcs Ts \u2212 j2\u03c6\u0307T 2 s ) \u02d9\u0303 \u03b4 + ( 1 \u2212 \u03c6\u03072T 2 s \u2212 j2\u03bcs Ts\u03c6\u0307 ) \u03b4\u0303 = \u2212 jKs \u03be\u0303 (19) Eq. (19) presents the differential equation of non-rolling frame between actuator deflection and angle of attack in complex form"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure12-1.png",
        "caption": "Fig. 12. Spherical body moving in the water\u2013vapor flow.",
        "texts": [],
        "surrounding_texts": [
            "The drag coefficient is calculated by a classical formula [7]: Cd \u00bc 2 Fd q A U2 ; \u00f05\u00de where Fd \u2013 the drag force, obtained from simulation results, q \u2013 water density, U \u2013 mean velocity of a water flow, m/s, A \u2013 reference area, m2."
        ]
    },
    {
        "image_filename": "designv11_13_0000782_j.ifacol.2015.06.433-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000782_j.ifacol.2015.06.433-Figure3-1.png",
        "caption": "Fig. 3. Fusion welding process",
        "texts": [
            " Considered as a joining method, this process or its variations are used for the deposition of material on a surface in order to recover worn parts or to form a coating with special characteristics. It was from the 19 th century that the welding technology has emerged on the world scenario contributing to the apperance of fusion welding processes with the discovery of acetylene by Edmund Davy, Sir Humphrey Davy experiences (1801- 1806) with the electrical arc and with the development of production sources of electricity [10]. The electrical arc used as a source of energy is the fusion welding process more used industrially Figure 3 represent a pioneer conception of this process. All additive manufacturing equipment sold today is based on the concept \"layer by layer\" and the material to be processed is what makes these devices different. However, the additive manufacturing using metal alloys is still restricted to a few components. The metal deposition in additive manufacturing technology may occur at different arc welding process (MIG, TIG and Plasma) Electro Beam Melting, Sective Laser Melting, Laser Cladding, Powders Sintering, among others [11]",
            " Considered as a joining method, this process or its variations are used for the deposition of material on a surface in order to recover worn parts or to form a coating with special characteristics. It was from the 19 th century that the welding technology has emerged on the world scenario contributing to the apperance of fusion welding processes with the discovery of acetylene by Edmund Davy, Sir Humphrey Davy experiences (1801- 1806) with the electrical arc and with the development of production sources of electricity [10]. The electrical arc used as a source of energy is the fusion welding process more used industrially Figure 3 represent a pioneer conception of this process. Fig. 3. Fusion welding process All additive manufacturing equipment sold today is based on the concept \"layer by layer\" and the material to be processed is what makes these devices different. However, the additive manufacturing using metal alloys is still restricted to a few components. The metal deposition in additive manufacturing technology may occur at different arc welding process (MIG, TIG and Plasma) Electro Beam Melting, Sective Laser Melting, Laser Cladding, Powders Sintering, among others [11]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure16-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure16-1.png",
        "caption": "Fig. 16. FEM results: deformation and contact normal stress.",
        "texts": [
            " The nodes either on the output end of the carrier shaft or on the outer circumference of the annulus gear are all defined as fixed. The Young\u2019s modulus and Poisson\u2019s ratio of all the components is equal to 206 GPa and 0.3, respectively. In order to compare with FEM, the distributed contact stress along the face-width and the load sharing among planets are calculated by the proposed LTCA approach and FEM. The torque with a value of 500 kNm is applied on the sun gear (input side). The FEM results are shown in Fig. 16 . Figs. 17 and 18 show the variation of the contact stress along the face-width of a sun-planet and an annulus-planet tooth pair, respectively. The results from the LTCA approach have a good agreement with those obtained from FEM, where the maximum difference is less than 10%. The variation of the contact stress in the annulus-planet tooth pair increases from the input side of the carrier to the output side, Fig. 17 , because of the larger stiffness of the carrier on the output side. This increasing trend of the contact stress is reduced in the sun-planet tooth pair and varies more evenly because of the larger twist stiffness of the sun gear on the input side, Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002312_1.4029054-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002312_1.4029054-Figure5-1.png",
        "caption": "Fig. 5 Model of impact process between outer race and swing supports",
        "texts": [
            " In addition, the generated impact force and the outer race and the supports\u2019 movements are different before and after the clearance elimination. Conditions with uneliminated and eliminated clearance would be separately discussed as follows. 3.3.1 Conditions With Uneliminated Clearance. If q(hs) 0, the clearance between the outer race and the supports hasn\u2019t been eliminated. According to the geometric analysis in Sec. 3.3, the model of impact process between the outer race and the swing supports can be simplified to that between the outer race and an equivalent clearance circle with a radius Rc(hs) (Rc(hs)\u00bcRo \u00fe q(hs)), as shown in Fig. 5. The support\u2019s reaction force acting on the outer race can be expressed: Fnx \u00bc Fn xoffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2 o \u00fe y2 o p \u00fe Ft yoffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2 o \u00fe y2 o p Fny \u00bc Fn yoffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2 o \u00fe y2 o p Ft xoffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2 o \u00fe y2 o p 8>>< >: (21) 062502-4 / Vol. 137, JUNE 2015 Transactions of the ASME Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www",
            " kn represents the contact stiffness between the outer race and supports; and dn is the penetration between the outer race and the support. dn can be expressed: dn \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2 o \u00fe y2 o q q\u00f0hs\u00de (23) and the rotational equation of motion for the outer race can be given: Io \u20acho \u00bc Mb1 FtRo (24) where Io represents the MOI of the outer race. Since these supports are separated and are only interconnected by a link, when any individual support swings, the rest ones would be driven to swing through same angles simultaneously. As shown in Fig. 5, both the normal impact force (Fn) and the frictional force in tangential direction (Ft) act on the impacted support. And the passive-driven supports are only acted by pull force (Fqj) from the link. Since those pull forces (Fqj) are very small, the contribution of pull forces (Fqj) are much smaller than that of impact force (Fn) when calculating the frictional forces in revolute joint 1. Thus, the influences of Fqj could be ignored. Based on the force analysis, the approximate total torque acting on all supports can be gained: Ms \u00bc Ftlt Fnln Fsrs (25) Therefore, the dynamic equation for supports\u2019 swing can be obtained: NsIs \u00fe lq Rl Il \u20achs \u00bc Ftlt Fnln Fsrs (26) where Ns is the number of the supports; Is and Il mean the MOI of each support and the MOI of the link, respectively; Rl is the radius of the link, as shown in Fig. 5; rs represents the radius of revolute joint 1; lq denotes the distance between the centers of revolute joint 1 and 2; ln and lt represent the force arms of Fn and Ft at the contact point A relative to revolute joint 1, as shown in Fig. 4. For concave support, ln and lt are ln \u00bc \u00f0Rm l\u00deOdOs OdOm sin\u00f0hs\u00de lt \u00bc Rm \u00f0Rm l\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 OdOs sin\u00f0hs\u00de OdOm 2 s 8>>< >>: (27) As to convex support, ln and lt can be gained: ln \u00bc \u00f0Rm l\u00deOdOs OdOm sin\u00f0hs\u00de lt \u00bc Rm \u00f0Rm l\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 OdOs sin\u00f0hs\u00de OdOm 2 s 8>>< >>: (28) According to Coulomb\u2019s law of friction [27], Fs can be gained: Fs \u00bc ld ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F2 n \u00fe F2 t p _hs > 0 ls ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F2 n \u00fe F2 t p _hs \u00bc 0 ld ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F2 n \u00fe F2 t p _hs < 0 8>>< >: (29) Generally, Ft\u00bc 0",
            " After the clearance eliminated, the radius of the equivalent clearance circle would be equal to or smaller than that of the outer race, namely, q(hs) 0 or Rc(hs) Ro. In this situation, if the displacement of the outer race is smaller than the value of {Ro Rc(hs)}, the outer race contacts with all of the supports. Otherwise, the outer race would only contact with some but not all supports. Conducted from these two situations, the elastic deformation for each support can be obtained: dnj \u00bc xo cos rj \u00fe yo sin rj q\u00f0hs\u00de (33) where rj is the angular position of each support, as shown in Fig. 5, rj \u00bc 2p\u00f0j 1\u00de\u00f0 \u00de=Ns\u00f0 \u00de 3=4\u00f0 \u00dep; j \u00bc 1; 2; Ns. In Eq. (33), if dnj< 0, the outer race has not contacted with the jth support. So, the value of dnj would be set as \u201c0\u201d under this circumstance. According to Hunt\u2013Crossley impact force law [26], the impact force between the outer race and each support can be expressed: Fnj \u00bc Nbknd 10 9 nj 1\u00fe 3 2 a _dnj (34) The components of each support\u2019s reaction force on x- and y-directions acting on the outer race can be written as Fnxj \u00bc Fnj cos\u00f0rj\u00de Fnyj \u00bc Fnj sin\u00f0rj\u00de (35) Journal of Engineering for Gas Turbines and Power JUNE 2015, Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001822_tro.2017.2671355-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001822_tro.2017.2671355-Figure6-1.png",
        "caption": "Fig. 6. Test 3: Initial and goal state of the cart\u2013pole. (a) Initial state. (b) Goal state.",
        "texts": [
            " This suggests that RHDDP allowed us to get a behavior that we could not achieve by using DDP, no matter the values of w1 , w2 , and wr . In this simulation, we tested RHDDP and DDP with a simple underactuated system: The cart\u2013pole (state dimension 4, control dimension 1). In order of priority, the tasks to perform were 1) keep the pole high (integral cost between 0.7N and N ); 2) reach a goal position with the cart (integral cost between 0.7N and N ). The two tasks are compatible only if we allow the system to use large control inputs (see Fig. Fig. 6 and first line of Table III). With DDP, we kept constant w1 = 1 (weight of the first task) and wr = 10\u22122 (weight of the regularization), whereas we varied w2 (weight of the second task). Using w2 = 10\u22121 or less, only task 1 was achieved (see corresponding lines in Table III) because wr was too large, with respect to w2 . Using w2 = 1 allowed the system to execute the second task, but it deteriorated the performance of the first task (see third line of Table III). With RHDDP, we used wr = 10\u22122 as well, but the hierarchy allowed for the execution of the second task, while not deteriorating the first task (see second line of Table III)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001250_j.matpr.2017.07.234-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001250_j.matpr.2017.07.234-Figure6-1.png",
        "caption": "Figure 6 CAD model of the turbine blade",
        "texts": [
            " For example, Argon gas with 5% to 20% CO2 is used for mild steel cladding, but for Aluminiumcladding 100% Argon gas is used. The wire diameter also controls the heat input and deposition resolution. Thinner wire is preferred for lesser heat input and better resolution of deposition. In this section, the proposed integrated substrate method has been used for the manufacturing of a turbine blade and the process is compared with the pure machining process. The manufacturing is done by following the steps given in Figure 5. The CAD model of a turbine blade is shown in Figure 6. First of all the build direction has to be identify for the deposition of the blade. One can observe from Figure 7 (a) (which shows the machining fixture of the blade) that the blade is fixed from both of the ends (root and shroud). The A-axis can move +45 to-90 and B-axis can move infinite rotation. Hence the orientation of the component for deposition can be in this range by considering the optimal machining. But for this study the orientation of the blade has been considered at A=0 and B=0 as shown in Figure 7 (b) and (c) to avoid any machining difficulty"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002066_2013.40949-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002066_2013.40949-Figure2-1.png",
        "caption": "FIG. 2 Cutaway view of a radial-ply tire.",
        "texts": [
            " It has four plies of cord fabric, each ply cut on the bias and each ply opposite in cord angle to the adjacent plies. The ply cords ex tend from bead to bead in a crisscross fashion to provide a strong, sound, long-life tire. This basic construction is used by tire manufacturers the world over. There may be two, four, six, eight or more plies, and in some cases a multiply bias-cut breaker in the crown area of the tire either on top of the plies or buried between plies to pro vide extra bruise strength where needed. Fig. 2 shows a similar cutaway of a radial-ply tire. It may be noted that its ply cords are not cut on the bias but instead are radial, running straight across from bead to bead. If we were to stop at this point, such a tire would be very unstable and very easy to dam age, as it would lack strength circumferentially. To improve stability and to give excellent strength and bruise re sistance in the crown area, a multiply Presented as Paper No. 59-111 at the Annual Meeting of the American Society of Agricultural Engineers at Ithaca, N"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003791_j.optlaseng.2019.01.009-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003791_j.optlaseng.2019.01.009-Figure2-1.png",
        "caption": "Fig. 2. Relationship between the coordinates of two plane reference systems.",
        "texts": [
            " In total there will be = \ud835\udc5b + 2 = 24 constraints or model functions: ( \ud835\udc9a , \u0302\ud835\udc99 ) = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239d \ud835\udc38 1 \u2212 \ud835\udc38 2 \ud835\udf081 \u2212 \ud835\udf082 [( \ud835\udf0e\ud835\udc65\ud835\udc65 ) 1 \u2212 ( \ud835\udf0e\ud835\udc67\ud835\udc67 ) 1 ]2 + 4 ( \ud835\udf0f\ud835\udc65\ud835\udc67 )2 1 \u2212 ( \ud835\udc41 1 \ud835\udc38 1 1 + \ud835\udf081 \ud835\udf06 \u210e\ud835\udc3e ) 2 [( \ud835\udf0e\ud835\udc65\ud835\udc65 ) 2 \u2212 ( \ud835\udf0e\ud835\udc67\ud835\udc67 ) 2 ]2 + 4 ( \ud835\udf0f\ud835\udc65\ud835\udc67 )2 2 \u2212 ( \ud835\udc41 2 \ud835\udc38 1 1 + \ud835\udf081 \ud835\udf06 \u210e\ud835\udc3e ) 2 \u22ef [( \ud835\udf0e\ud835\udc65\ud835\udc65 ) 22 \u2212 ( \ud835\udf0e\ud835\udc67\ud835\udc67 ) 22 ]2 + 4 ( \ud835\udf0f\ud835\udc65\ud835\udc67 )2 22 \u2212 ( \ud835\udc41 22 \ud835\udc38 1 1 + \ud835\udf081 \ud835\udf06 \u210e\ud835\udc3e ) 2 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 0 0 0 \u22ef 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 (21) In which, for simplicity, ( \ud835\udf0exx ) i , ( \ud835\udf0ezz ) i and ( \ud835\udf0fxz ) i have not been relaced by their expressions from Eqs. (1) to (4) . In these latter expresions the coordinates x and z must be calculated as those of the origin, 0 and z 0 , plus those relative to them, the measured coordinates x\u2019 and \u2019 ( Fig. 2 ): = \ud835\udc65 0 + \ud835\udc65 \u2032cos \ud835\udf03 + \ud835\udc67 \u2032sin \ud835\udf03 = \ud835\udc67 0 + \ud835\udc67 \u2032cos \ud835\udf03 \u2212 \ud835\udc65 \u2032sin \ud835\udf03 (22) In order to minimize the objective function defined in Eq. (17) , the ethod of Lagrange multipliers defines an auxiliary function which in- orporates the constraints given by Eq. (21) into Eq. (17) : ( \ud835\udc9a , \u0302\ud835\udc99 , \ud835\udf40; \ud835\udc99 ) = ( \ud835\udc99 \u2212 ?\u0302? ) \ud835\udc47 \ud835\udeba\u22121 ( \ud835\udc99 \u2212 ?\u0302? ) + 2 \ud835\udf40\ud835\udc47 \ud835\udc21 ( \ud835\udc9a , \u0302\ud835\udc99 ) (23) w p i \u2207 w d i t \ud835\udc99 t f \u239b\u239c\u239c\u239c\u239c\u239d t p d ( m a \ud835\udc62 \ud835\udc5f w t fi 4 s s d a t t T b e a f \ud835\udf08 m o fi s fi r c m s t S n here \ud835\udf40 = ( \ud835\udf061 , ..., \ud835\udf06\ud835\udc51 ) \ud835\udc47 are a set of parameters known as Lagrange multiliers [22] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002479_j.ifacol.2015.11.274-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002479_j.ifacol.2015.11.274-Figure4-1.png",
        "caption": "Fig. 4. Sample trajectory. Regional control laws can be used in the gray and light blue regions.",
        "texts": [
            " In all these cases, |W| = 1 and the border is defined by the additional condition \u03bb\u22c6 W = 0. We were able to compute these borders with pseudo arc length parameter continuation, but do not give details here. We note that we did not calculate the domain of stability. Additional regions in which the MPC control law that results for the particular choice of P , Q, R and N is known not to be stabilizing have been omitted in Figs. 2 and 3. Finally, we illustrate the use of the implicitly defined regional control laws in Algorithm 1. Figure 4 shows the trajectory of the closed-loop system that results for x0 = (2, 0.0105)\u2032. In the gray and light blue regions shown in Fig. 4, the control law (16) can be reused instead of solving the MPC problem. The optimal input signals that resulted from the control law (16) were equal to those from solving the MPC problem (5) within the numerical tolerance of the solvers. We carried out this comparison for about 4000 initial conditions in the gray and light blue regions in Fig. 4. We stated conditions under which a local solution to a nonlinear MPC problem characterizes a regional optimal feedback law. While we illustrated our results with a numerical example, the purpose of this paper actually was to explore the idea of regional optimal feedback in nonlinear MPC for the first time. We point out several directions for future research. We mentioned in Section 3 that it must be investigated whether there exist special problem classes for which the regional optimal feedback law (16) can actually be found analytically (or described in an otherwise convenient way)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001493_tdei.2018.006879-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001493_tdei.2018.006879-Figure2-1.png",
        "caption": "Figure 2. Particle trajectory comparison of test and simulation.",
        "texts": [
            " In a uniform electric field, the particles (neutral in charge) will be affected by an electric field and particles will have a directional movement in accordance with the direction of the electric field orientation, or orientation polarization force. The particle polarization process is shown in Figure 1. In a non-uniform electric field, under the action of electric field force, the metal particles also can achieve a state of balance. However, unlike strong electric field, the energy for the achievement of balance for particles is relevant to their position. The particle polarization in non-uniform electric field is shown as in Figure 2. According to electrostatic field theory, when the metal particle moves in an electric field E, the electric field boundary conditions on particle surface can be expressed as [30]: Constant  (1) 1r      (2) and the space potential solutions can be determined from [31]: 3 2 cos cos Er ER R      (3) Here R is the distance between the calculating point and the center of sphere. Suppose placing electric dipole P in its center, the produced potential at R can be derived from the Equation (4): 2 1 1 cos 4 4 P P R R       g (4) From (3) and (4), we know that the induction electric dipole in the electrostatic field can be derived from the Equation (5): 34P r E (5) In order to facilitate analysis, somewhere near particles we conducted infinitesimal analysis to the hypothesis particles (E1) and (E2) whose positions are indicated in the figure below, between the motion of the dipole of P1 and P2, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure13-1.png",
        "caption": "Fig. 13. Von Mises and deformations with elastic housing for F r = 6 kN and \u03c9 = 500 rad/s at (a,b) t = 1 \u00d7 10 \u22123 s (c,d) t = 2 \u00d7 10 \u22123 s (e,f) t = 3 \u00d7 10 \u22123 s.",
        "texts": [
            " 12 , the Von Mises stress field in dynamic mode recalls the profiles of the load distribution observed in Fig. 11 (a). The maximum stress stays close to \u03c8 = 0 o and follows the contact displacement. Mechanical forces at the discretized interface are transmitted within rings resulting in noisy sollicitations. The stress magnitude is determined between 200 MPa and 300 MPa according to time. Due to the rotation, the outer ring feels dynamic loading. In multimedia component 2 given in Appendix A or snapshots Fig. 13 (a,c,e) show that deformations of rings lead to a more active stress zone than rigid housing. The maximum is still observed near \u03c8 = \u00b140 o as described in the static case and follows the contact position. The magnitude oscillates between 300 MPa and 400 MPa, typically observed on the outside diameter of the outer ring. Furthermore, analysis of the deformation of the outer ring considering half-clamping is given in Fig. 13 (b,d,f) where the displacement is amplified by a factor 50. The radial load ovalizes the ring and the movement of rolling elements induces deformation modes in dynamic. Displacement values do not exceed 20 microns for F r = 6 kN. At the same instant, a correlation is observed between the load distribution, stresses and deformations. This kind of observation validates the assumption of a non-constant and anisotropic clearance. These deformations also exist in the case of rigid housing but they remain confined at contacts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000169_ext.12016-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000169_ext.12016-Figure1-1.png",
        "caption": "Figure 1 Round rolling process with two tools.",
        "texts": [
            " The most common methods are still cutting techniques, which made a continuous progress during the last century. In the past few years, it has been shown that round rolling also has capabilities to manufacture gearings with higher number of teeth.1 Round rolling is an efficient incremental coldmassive forming process. It is already used successfully either to finish prefabricated high gears or to form gearings with lesser number of teeth. In this study, we are focused on forming complete high gears into a full material. Figure 1 illustrates this process in three steps. At first, the component is clamped between the tips. This mounting allows the rotation and an axial motion of the component. Then, the tool is fed to the initial diameter. The start of the tool rotation and an initial penetration to roll the exact number of teeth follows as second phase. In the third phase, both tools penetrate into the component until the desired tooth root diameter is reached. A balanced material flow along the tool teeth and a nearly concentric axial component position are managed by several changes in the rolling direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002587_tmag.2014.2325671-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002587_tmag.2014.2325671-Figure6-1.png",
        "caption": "Fig. 6. CoPt microcylinder rolling on a substrate under a rotating magnetic field.",
        "texts": [
            " The experiments were performed with CoPt microtubes without etching away the mandrel to increase rigidity and reduce risk of breakage during handling. This was not needed for the CoNi cylinders as they have relatively thicker walls. The soft-magnetic CoNi aligns as expected its easy axis (long axis) with the field, whereas the hard-magnetic CoPt aligns its axis perpendicular to the field due to its set magnetization direction. The CoPt microcylinders can be rotated about their axis if a rotating magnetic field is applied and this motion can be used to steer the microcylinder by rolling on the surface. Fig. 6 shows the motion of the microcylinder as it is steered on a substrate. The Co-based hard- and soft-magnetic microtubes were grown by electrodeposition on cylindrical substrates. Hardmagnetic properties were achieved without any postprocessing techniques, (e.g., annealing) with CoPt alloys. Thick structures (up to 10 \u00b5m) were demonstrated maintaining the magnetic characteristics. Soft-magnetic properties were obtained with coercivities below 1.6 kA/m (20 Oe) using CoNi alloys. The magnetic properties of both materials can be maintained when plating on long segmented wires, thus enabling batch fabrication of micrometer-sized cylinders"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003973_icieam.2019.8743076-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003973_icieam.2019.8743076-Figure1-1.png",
        "caption": "Fig. 1. 2D motor geometry.",
        "texts": [
            " It is relevant to fully understand how inter-turn shortcircuit faults of different severity levels affect the temperature of induction machines. Accordingly, next sections provide a detailed explanation on the development and implementation the finite elements method (FEM) to the modeling of an induction motor while subjected to inter-turn short-circuit faults. II. FINITE ELEMENTS MODEL The FEM model of the induction motor under study was developed resorting to the software tool Flux2D (Cedrat). Fig. 1 depicts the 2D geometry of the motor, used in Flux2D software. Each motor element is properly labelled in the figure. TABLE I provides the most relevant technical parameters of the motor considered for this study. To obtain the map of the temperature distribution inside the motor, a two-stage procedure is adopted. The electromagnetic behaviour is initially simulated to acquire data of the losses occurring in each element of the motor. Iron losses, Joule losses and additional losses are computed through the electromagnetic simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000599_j.engfailanal.2012.02.008-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000599_j.engfailanal.2012.02.008-Figure10-1.png",
        "caption": "Fig. 10. Detail of field of equivalent stresses on pedestal traverse.",
        "texts": [],
        "surrounding_texts": [
            "Computational model of traverse (due to symmetry only one half of beam and middle part with arm of toothing was considered) is given in Fig. 7. Computation of traverse was accomplished for all load cases given above. In Fig. 8 is given field of equivalent stresses on traverse of casting pedestal for the most danger state of loading. In Figs. 9 and 10 are shown details of equivalent stresses on upper side of traverse beam."
        ]
    },
    {
        "image_filename": "designv11_13_0000097_educon.2012.6201108-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000097_educon.2012.6201108-Figure6-1.png",
        "caption": "Fig. 6. DC motor-generator and Arduino top view (development stage)",
        "texts": [
            " Two servo motors, sensor board and video camera Sensor board Video camera 2nd servo motor axle line 1st servo motor axle line SN74LV4066). Switches also are controlled from Arduino board. Voltage dividers are used to fit in to ADC input upper level +5V. Low pass filters on all ADC inputs eliminate PWM fluctuations and allow reading average values. LCD display is driven from Arduino board via 4-wire connection. Another Arduino board in connection with Arduino Ethernet board provides connection to internet (Fig. 6.) by means of standard TCP/IP protocol. Both Arduino boards communicate via standard serial TTL connection (19200 bps baud rate). Voltage and current ADC readings are sent to Arduino Ethernet board through serial communication. Atmega328 microcontroller (second Arduino board) controls Allegro A3967 micro-stepping driver with translator, loaded by 2-phase bipolar step motor. A3967 driver allows motor axle rotation with full step, 1/2 step, 1/4 step and 1/8 step movement. Micro-stepping is achieved by motor phase control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001848_j.procir.2015.03.095-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001848_j.procir.2015.03.095-Figure7-1.png",
        "caption": "Fig. 7. (a) 3D reality representation created from the dexel model, (b) photo realistic visualization of the simulated gear.",
        "texts": [
            " For the simulation of the cutting tool geometry a function f(x) has to be given (Fig.5a). The full hob is created by a helical transformation of the tooth profile into appropriate positions, gaining 420 teeth ordered in two numbers of starts. Fig. (5a) shows the cutting profile and a simulated gap of dexel disk (Fig.5b). Depending on the kinematical motion not all teeth are equally used. Fig. 6 shows the statistics analysis of the removed material and the cutting teeth in use for a singe of dexel disk. The mostly used cutting edges are marked in Fig.5a. In Fig. 7a the gear shape is visualized as a VRML (Virtual Reality Modeling Language) model. For a better visualization and verification all dexel disks in VRML visualization are separated with some free space. The geometry can be exported at any process step. A high quality photo realistic visualization is enabled due to an integrated ray-tracing export (Fig.7b). The calculation time of the dexel simulation data (results Fig. 5 & 6) by model (1) and workpiece discretization mentioned above (section 2) is just 50 seconds by Intel i53210M without additional mathematical accelerations. The high quality image with a resolution of 2000 points for one dexel disk has a calculation time of 277 seconds. As it is mentioned before, the quaternion algebra together with the dexel approach enables an efficient numerical simulation of the full gear hobbing process. Nevertheless the dexel model itself cannot predict thermal deformation that\u2019s why an external solver is used additionally"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002532_tasc.2016.2526025-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002532_tasc.2016.2526025-Figure1-1.png",
        "caption": "Fig. 1. Cross section of the Maglev system.",
        "texts": [
            "ndex Terms\u2014Damping, electrodynamic suspension, quenching. I. INTRODUCTION THE superconducting magnetically levitated train (JR Maglev) has been developed and the commercial line (from Tokyo to Nagoya) has been under construction [1]. Fig. 1 shows picture of the Maglev system [2]. The null-flux eight-figure coils for levitation and guidance are set on the side wall of the guideway. The armature winding of the linear synchronous motor (LSM) is put outside of the levitation coils. The bogie on the vehicle has the superconducting (SC) coils. These coils are used not only for the levitation magnet of the EDS but the field magnet of the LSM [3]. Levitation and guidance force are generated by passing the SC coil in front of the levitation coils",
            " The electric circuit equations for the levitation and guidance coil are given as follows: \u2212\u2202\u03c6c1 \u2202t + \u2202\u03c6c2 \u2202t \u2212 ic,1R+ ic,2R =0 (1) \u2212\u2202\u03c6c3 \u2202t + \u2202\u03c6c4 \u2202t \u2212 ic,3R+ ic,4R =0 (2) 1051-8223 \u00a9 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. where ic,k (k = 1, 2, 3, 4) is the kth levitation coil current. R is levitation coil resistance. The subscript 1\u20134 means the coil number as shown in Fig. 1. The one side of the upper coil is Coil 1 and the lower Coil 2. And the other side the upper coil is Coil 3 and the lower Coil 4. All the levitation coils consist of this four pair configuration. \u03c6ck (k = 1, 2, 3, 4) is linkage flux of the kth levitation coil. It includes the flux from the SC coils and also the levitation coils equipped around the coil. The null flux connection is described as follows: \u2202\u03c6c1 \u2202t + \u2202\u03c6c3 \u2202t + ic,1R\u2212 ic,3R+ 2Rn(ic,1 + ic,2) = 0. (3) Rn is the resistance of the null flux connection cable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003460_s12206-018-0441-0-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003460_s12206-018-0441-0-Figure9-1.png",
        "caption": "Fig. 9. New definition of the Sub-problems for inverse kinematics.",
        "texts": [
            " r r e st e p e e e e e g p G p= - (25) In this example, pr is a random point that does not exist in S6 and G3(\u03b8) is transform matrix between end-effector pose and last rotation angle. Subproblems 2 and 3 are not intuitive for a robot beginner, and the two skew axis vectors, such as those shown in Fig. 8, are not solved because two screw axis are not intersection. In this section, a new subproblem is introduced to solve the inverse kinematics when the two axis vectors are skewed in space with a zero pitch and a unit magnitude. New subproblems include the subproblems 2 and 3, as shown in Fig. 9. A new subproblem provides a solution of inverse kinematics for all cases in space and is geometrically explained using skew coordinates. 4.1 Introduction of skew coordinates In this section, skew coordinates are introduced to define a new subproblem. Skew coordinates are useful to explain the physical meaning of two skew axis vectors. Therefore, the skew coordinates are more complicated than the Cartesian coordinates because the two axis vectors (S1, S2) are not orthogonal to each other (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002953_1045389x16679018-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002953_1045389x16679018-Figure1-1.png",
        "caption": "Figure 1. Plate geometry and notation",
        "texts": [
            " In this case, it is possible to write ek p =Dpu k , ek n =Dnu k \u00f04\u00de defining Dp = \u2202 \u2202x 0 0 0 \u2202 \u2202y 0 \u2202 \u2202y \u2202 \u2202x 0 2 64 3 75 \u00f05\u00de Dn = \u2202 \u2202z 0 \u2202 \u2202x 0 \u2202 \u2202z \u2202 \u2202y 0 0 \u2202 \u2202z 2 64 3 75= 0 0 \u2202 \u2202x 0 0 \u2202 \u2202y 0 0 0 2 64 3 75 zfflfflfflfflfflfflfflfflffl}|fflfflfflfflfflfflfflfflffl{DnO + \u2202 \u2202z 0 0 0 \u2202 \u2202z 0 0 0 \u2202 \u2202z 2 64 3 75 zfflfflfflfflfflfflfflfflffl}|fflfflfflfflfflfflfflfflffl{Dnz \u00f06\u00de The constitutive relations and variational statements for each problem are defined in the following sections. Stress components for a generic layer k can be obtained by means of the Hooke\u2019s law sk =Ckek \u00f07\u00de The elastic coefficients of the matrix C are expressed in the problem reference system (i.e. (x, y, z) system reported in Figure 1). The dependence of the elastic coefficients Cij on Young\u2019s modulus, Poisson\u2019s ratio, the shear modulus and the fibre angle is not reported. A detailed discussion is reported in the book by Reddy (1997). The stress components can be grouped into in- plane (p) and out-of-plane (n) components as the strain components, i.e. sk p = sk xx sk yy sk xy h iT , sk n = sk xz sk yz sk zz h iT \u00f08\u00de In this case, the Hooke\u2019s law can be defined as sk p =Cppek p +Cpnek n sk n =Cnpek p +Cnnek n \u00f09\u00de In the case of anisotropic materials it is possible to write Ck pp = Ck 11 Ck 12 Ck 16 Ck 12 Ck 22 Ck 26 Ck 16 Ck 26 Ck 66 2 64 3 75, Ck nn = Ck 55 Ck 45 0 Ck 45 Ck 44 0 0 0 Ck 33 2 64 3 75, Ck pn =CkT np = 0 0 Ck 13 0 0 Ck 23 0 0 Ck 36 2 64 3 75 \u00f010\u00de The analysis of a plate can be conducted by means of the principle of virtual displacements (PVD), that states XNL k = 1 dLk int = XNL k = 1 dLk ext \u00f011\u00de Here d denotes the virtual variation, dLk int is the virtual variation of the strain energy that is computed considering the internal stresses and strain distributions for a generic k layer, dLk ext is the virtual variation of the work made by the external loadings on the generic layer k and NL is the total number of layers of a multilayered plate",
            " In the case of pure-mechanical analysis, the virtual variation of the strain energy can be computed as dLk int = XNL k = 1 Z Vk dek p sk p + dek n sk n dVk = XNL k = 1 Z Ok Z Ak dek p sk p + dek n sk n dOk dzk \u00f012\u00de where Vk is the volume of the layer k. The virtual variation of the external loading can be computed as dLk ext = XNL k = 1 Z Vk duTk pk dVk \u00f013\u00de where duk is the virtual variation of the displacement vector uk and the components of the vector pk are the load distributions according to the reference system axes x, y and z. The operator R Ak dzk denotes the integration along the thickness direction. Here Ok is the reference surface of the generic k layer (see Figure 1). In this work, the uncoupled thermo-mechanical analysis is performed in which the temperature is considered as an external load. A plate subjected to a temperature distribution can be analysed by defining the thermal stresses as sk pT =Ck pp ek pT +Ck pn ek nT sk nT =Ck pn ek pT +Ck nn ek nT \u00f014\u00de where epT and enT indicate the in-plane (p) and out-ofplane (n) strains due to a temperature gradient, that is ek pT = ek xxT ek yyT ek xyT h i , ek nT = ek xzT ek yzT ek zzT h i \u00f015\u00de Considering the thermal expansion coefficient vector a, it is possible to write ek pT = ak 1, ak 2, 0 uk(x, y, z)=ak p uk(x, y, z) ek nT = 0, 0, ak 3 uk(x, y, z)=ak n uk(x, y, z) \u00f016\u00de where uk(x, y, z) is the relative temperature distribution in a generic k layer referred to a reference temperature ue",
            " In this case, the matrix of differential operators can be expressed as De = \u2202, x 0 0 0 \u2202, y 0 0 0 \u2202, z 2 4 3 5= \u2202, x 0 0 0 \u2202, y 0 0 0 0 2 4 3 5 |fflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl} DeO + 0 0 0 0 0 0 0 0 \u2202, z 2 4 3 5 |fflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflffl} Dez \u00f028\u00de The analysis of a piezoelectric plate can be conducted by means of the PVD (see (11)) extended to the electro-mechanical problem, that states XNL k = 1 Z Vk dekT p sk p + dekT n sk n dEkT ~D k dVk = XNL k = 1 dLk ext \u00f029\u00de Plate geometry is reported in Figure 1, the reference surface is denoted as O and its boundary as G. The reference system axes which belong to the reference surface O are denoted as x, y and z is the thickness coordinate. The length side dimensions of the plate are indicated as a and b, and the thickness of the plate is defined as h. In the framework of the CUF, the displacement field of a plate can be described as u(x, y, z)=Ft(z) ut(x, y), t = 1, 2, . . . ,N + 1 \u00f030\u00de where a summation on the index t is implied according to the Einstein notation. Here u is the displacement vector (ux uy uz), whose components are the displacements along the x, y, z reference axes (see Figure 1), Ft are the so-called thickness functions depending on z and ut =(utx, uty, utz) are the displacement variables depending on the in-plane coordinates x, y; N is the order of expansion. The expansion functions Ft can be defined on the overall thickness of the plate or for each k-layer. In the former case equivalent single layer (ESL) approach is followed, while, in the latter case, a layer-wise (LW) approach is used. Examples of ESL and LW schemes are reported in Figures 3(a) and 3(b), respectively: a transverse section of a multilayered plate is reported, the number of layers is equal to NL"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001432_s00170-018-1990-1-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001432_s00170-018-1990-1-Figure4-1.png",
        "caption": "Fig. 4 The FE model with mesh and boundary conditions",
        "texts": [
            " The minimum dimension ratio of the sheet and rivet is larger than 6, which is supposed to be enough for a negligible effect of the stress free out boundaries [22]. The clearance between the hole wall and rivet shank is 0.15 mm. Material parameters for the 2117-T4 Al alloy slug rivet and 7050-T7451 Al alloy sheet are summarized in Tables 1 and 2, where \u03b5 is the true strain and \u03c3 is the true stress [24, 25]. The FE model is generated using ABAQUS 6.14 with CAX4R reduced integration 4-node axisymmetric elements (Fig. 4). Three deformable bodies, two sheets and a rivet, are defined in the model, while the riveting dies and pressure feet are defined as rigid bodies. A typical mesh generation technology is applied [26]. Mesh size of rivet is 0.06 mm. Mesh size of the area in the vicinity of the rivet hole is 0.08 mm. And, mesh size of the area far from rivet hole is 1.2 mm. The sheet surface on the far-end is constrained in X-direction with the Y-direction nodes constrained at the top and bottom [24]. All freedom degrees of the riveting dies and pressure feet are constrained, except the Y-direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000195_tmag.2013.2245878-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000195_tmag.2013.2245878-Figure8-1.png",
        "caption": "Fig. 8. M1: Spindle motor with 12 slots and 5 magnetic pole-pairs.",
        "texts": [
            " \u2022 Deduction-1: IUMP has a constant component in eccentricity direction only. \u2022 Deduction-2: In one motor revolution, the lowest order of the IUMP harmonic is two times that of its pole-pair number. \u2022 Deduction-3: The orders of the IUMP harmonics are independent of the eccentricity. For verifying the analytical results obtained in Section II, a misalignment PMSM motor is simulated with a 3-D finite element method (FEM), and the FEM results are compared with the analytical ones. The PMSM used has 12 stator slots and 5 magnetic pole-pairs, and its structure is shown in Fig. 8. The motor will be called M1 in the paper. M1 has the rotor with surfaced mounted magnet. Its detailed dimensions are shown in Table I. In the simulation, the motor is processed with ten segments as illustrated in Fig. 9. The eccentricity ratio and the direction of eccentricity are thus different in each segment. Therefore, the IUMP should be different in each segment and the IUMP center may vary in different segments. In order to calculate the IUMP accurately, the stator is processed to be formed by ten segments in axial direction, and four sections in the plane, and they are named Sections A, B, C, and D, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003312_icsens.2017.8234170-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003312_icsens.2017.8234170-Figure2-1.png",
        "caption": "Fig. 2. Configuration of proposed novel sensor device.",
        "texts": [
            " Therefore, in this study, we propose a novel sensor device that can directly observe the biological information (pH and EC). We also fabricated a prototype of the proposed device and evaluated it through experiments with mimicking plants. II. SENSOR DESIGN In this section, to quantitatively monitor plant health conditions, we propose a novel sensor device to obtain the pH values of phloem sap containing nutritional substances produced by photosynthesis and the EC values a xylem sap containing minerals. Furthermore, the configuration and various functions mounted on the proposed device are described. Figure 2 shows the configuration of the proposed novel sensor device mounting with pH and EC measurement functions and a conceptual application on actual plants. The sensor device consists of three Si cantilevers and a pedestal. In detail, the device mounts the following three functions (1) the ISFETbased pH sensor (ISFET and reference electrode), (2) the EC 978-1-5090-1012-7/17/$31.00 \u00a92017 IEEE sensor, and (3) the p-n diode temperature sensor. It is possible to acquire pH and EC information by inserting the sensor device from the vertical direction into the plant stem"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000735_icra.2014.6907448-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000735_icra.2014.6907448-Figure1-1.png",
        "caption": "Fig. 1. Diagram of a sagging cable",
        "texts": [
            " Section IV presents an example of a 6-DOF cable-suspended manipulator, and discusses the effects of dynamic sagging cable by numerical simulations and experiments. The effects of cable sag on the static pose error are also experimentally investigated. An explicit comparison with the results of other methods available in literature is presented. Finally, conclusions are made in Section V. 978-1-4799-3685-4/14/$31.00 \u00a92014 IEEE 4055 In this section, improved dynamic stiffness matrix of a single cable is formulated by considering the cable mass, elasticity and damping [29]. The model for an inclined cable is presented in Fig. 1. One of the cable-ends is fixed, and an external force is applied to the other end of the cable. Under the effect of external force and gravity, the shape of the cable between points A and B is not a straight line, but a sagging curve. The cable is considered as a continuum and its shape is given by l (the chord length), d (the sag perpendicular to the chord), and \u03b1 (the inclination angle). According to [29], the following assumptions are made. \u2022 The cable is assumed to be uniform with unstrained cross section area A, mass per unit length m, and linear Young\u2019s modulus E"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002641_j.ijepes.2016.04.050-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002641_j.ijepes.2016.04.050-Figure2-1.png",
        "caption": "Fig. 2. The BDFRG: (a) schematic cross section [13], (b) axially-laminated reluctance rotor [13], (c) wound stator core [14], and (d) wound stator core in the frame.[15]",
        "texts": [
            " Also, elimination of brushes and slip rings in its structure ensures high reliability and lowmaintenance cost of the BDFRG in comparison with DFIG, which are especially important to off-shore plants [12]. The BDFRG has two sinusoidal distributed three-phase windings in its stator with different pole pairs and supply frequencies. A reluctance rotor with Pr salient poles can make a magnetic coupling between the primary (with P1 pole pairs) and secondary (with P2 pole pairs) [2]. The primary is directly connected to the grid while the secondary is connected through a back to back converter. The BDFRG connection to the grid is shown in Fig. 1 [12], and Fig. 2 illustrates its structure. In literature, different methods have been proposed to control of the BDFRG under balanced condition. These methods can be classified into following categories: scalar control [16\u201319], field orientation control (FOC) [20\u201328], direct torque control (DTC) [23,29\u201334] and direct power control (DPC) [12,35]. A comparative analysis of these control methods can be found in [36]. All of these methods are designed to control of the BDFRG under balanced condition, whereas, wind power plants are often installed in rural and remote areas which weak grid with unbalanced voltages is usual [37\u201339]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003766_rcs.1983-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003766_rcs.1983-Figure3-1.png",
        "caption": "FIGURE 3 Coordinate system of the joint. Ob,Od , and Og denote the reference coordinate and the base of the joint, the tip of the joint, and the tip of the grasper, respectively",
        "texts": [
            " The electric and pneumatic components in the control box are the same as those for the joint actuation mechanism. Each subscript i (1-4 and g) of Fi, Pij, Aij, and ui is correspond the wire holes in Figures 1 and 2. Aia and Aib represent the cylinder's chamber of the rod side and the nonrod side, respectively. In this paper, these cylinder forces and the input voltages are set by the PID controllers. The proposed estimation method is based on a model of the wrist joint and the measured position and driving force. In Figure 3, the coordinates Ob,Od , and Og denote the reference coordinate and base of the joint, the tip of the joint, and the tip of the grasper, respectively. \ud835\udcc1d and \ud835\udcc1g represent the vector from Ob to Od and the vector from Od to Og . The tip of the grasper Og is coresspond the tool tip p. First, we discuss the kinematics of the wrist joint by using a 2-DOF tendon-driven continuum model.30 From the symmetry of the antagonistic drive, the relationship between each tendon-driven joint position \ud835\udf19 = [\ud835\udf19y, \ud835\udf19x]T and cylinder position Xi can be obtained geometrically as follows: \ud835\udf19(X) = [ \ud835\udf19y \ud835\udf19x ] = \u2212 1 2r [ X1 \u2212 X3 X2 \u2212 X4 ] , (1) r shows the distance between the center of the joint and the wire hole"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003410_j.matpr.2017.11.530-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003410_j.matpr.2017.11.530-Figure4-1.png",
        "caption": "Fig. 4: Simulation model and basic cutting parameters",
        "texts": [
            " In recent years, commercial FE packages such as DEFORM 2D/3D, ABAQUS, ADVANTEDGE, LS-DYNA etc., It has been used excessively in both academic and industrial world for process analysis [8]. DEFORM (Design Environment for Forming) is a Finite Element Method based system that can be applied to several manufacturing processes such as forging, rolling, forming, heat treatment process and machining. Deform has a specific machining module to quickly set up turning, milling, boring and drilling operations. The figure 4 shows that the simulation of process with varying parameters such as speed, feed, depth of cut etc. The choice of finite element software is extremely important to determine the quality and accuracy of the analysis. In this study, DEFORM 3D was used (this software has been exclusively built with machining operations in mind, thus being optimized for metal cutting operations). It has a user friendly interface and can simulate both 2D and 3D machining operations with a wide range of materials in the property library"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003621_j.apm.2018.10.002-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003621_j.apm.2018.10.002-Figure2-1.png",
        "caption": "Figure 2: General sketch for the new screw-based sub-problem 2",
        "texts": [
            " Based on two simple geometrical evidences, the following equation set can be obtained for the new screw-based sub-problem 1:    \u2016qN1 \u2212 b\u2016 = \u2016qN1 \u2212 cN2\u2016, (b\u2212 cN2)\u03c9T N1 = 0 \u2016qN2 \u2212 cN2\u2016 = \u2016qN2 \u2212 cN1\u2016, (cN2 \u2212 cN1)\u03c9T N2 = 0 \u2016qN3 \u2212 cN1\u2016 = \u2016qN3 \u2212 a\u2016, (cN1 \u2212 a)\u03c9T N3 = 0 (7) where the symbol \u201d\u2016 \u00b7 \u2016\u201d denotes the calculation of the length of one vector, and \u03c9N1, \u03c9N2, and \u03c9N3 are unit rotational vectors in the axial directions of joints N1, N2, and N3 in the robotic base frame, respectively. After finding cN1 and cN2, the problem can be converted into three PK-1 sub-problems, which can be expressed as:    exp(\u03be\u0302N1\u03b8N1)c \u2032 N2 = b \u2032 exp(\u03be\u0302N2\u03b8N2)c \u2032 N1 = c \u2032 N2 exp(\u03be\u0302N3\u03b8N3)a \u2032 = c \u2032 N1 (8) Therefore, the three angular displacements in Figure 1 can be solved based on the solution to the PK-1 sub-problem described in literature report [24]: \u03b8m = arctan 2(\u03c9m(u\u2217 m \u00d7 v\u2217m)T ,u\u2217 mv \u2217 m T ),m = N1, N2, N3 (9) ACCEPTED M The general sketch of the new screw-based sub-problem 2 is presented in Figure 2. This time, the inverse motion can be decomposed into four rotational motions: rotation about joint N1 from b to cN3, rotation about joint N2 from cN3 to cN2, rotation about joint N3 from cN2 to cN1, and rotation about joint N4 from cN1 to a. The forward description of the inverse motion from b to a can be given as: { exp(\u03be\u0302N1\u03b8N1)RN2 exp(\u03be\u0302N3\u03b8N3) exp(\u03be\u0302N4\u03b8N4)a \u2032 = b \u2032 , if RN2 is given exp(\u03be\u0302N1\u03b8N1) exp(\u03be\u0302N2\u03b8N2)RN3 exp(\u03be\u0302N4\u03b8N4)a \u2032 = b \u2032 , if RN3 is given (10) where RN2 and RN3 present the rotational motions about axes N2 and N3, respectively, and for the robotic IK problems they are equal to a single joint motion or the sum of two or more joint motions. Based on Figure 2, the rotational screw motions can be expressed as RN2 = exp(\u03be\u0302N2\u03b8N2) and RN3 = exp(\u03be\u0302N3\u03b8N3). According to the geometric relationships (Figure 2), when RN2 is given a set of equations can be used to determine cN1 = (xN1, yN1, zN1), cN2 = (xN2, yN2, zN2), and cN3 = (xN3, yN3, zN3), given as:    RN2c \u2032 N2 = c \u2032 N3, \u2016qN1 \u2212 b\u2016 = \u2016qN1 \u2212 cN3\u2016, (b\u2212 cN3)\u03c9T N1 = 0, \u2016qN3 \u2212 cN2\u2016 = \u2016qN3 \u2212 cN1\u2016, (cN2 \u2212 cN1)\u03c9T N3 = 0, \u2016qN4 \u2212 cN1\u2016 = \u2016qN4 \u2212 a\u2016, (cN1 \u2212 a)\u03c9T N4 = 0, (11) where qN1, qN2, and qN3 are the points located on axes N1, N2, and N3, respectively, and for the computation they are selected based on the relative ease of solving the equation set. By solving equation set (11), position vectors of points cN1, cN2, and cN3 can be determined"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000507_1.4711923-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000507_1.4711923-Figure5-1.png",
        "caption": "FIG. 5. Light-driven dynamic plasmonic roll-up. (a, b) Dynamic control of plasmonic roll through broadband light irradiation. (a) Incident light rolls up film, and (b) after switching off light, it unrolls. (c, d) Video stills showing starting position of a plasmonic roll (c, white dashed line) and displaced position after 1.5 s of light irradiation (d, red dashed line) (enhanced online) [URL: http://dx.doi.org/10.1063/1.4711923.1].",
        "texts": [
            " 4(c)), clearly show the hexagonally patterned regular array of nanopore holes and the PDMS layer (Fig. 4(i)). These rolls show colours strongly dependent on the exact nanostructure morphology. The plasmonic tubes are loosely rolled multilayers (Fig. 4(h)), though single layer tubes can be fabricated by careful control of parameters.13 Such non-bonded multilayers however allow dynamical tuning of the rolls. The plasmonic tubes are found to be very sensitive to light irradiation. Upon broadband illumination, they quickly roll inward (Fig. 5(a)). This is because optical irradiation heats up the tubes and as a consequence of the elevated temperature, the gold films expand and PDMS films shrink. These two effects drive the tubes to roll inward (as the gold film is outside and the PDMS is inside the double-layer film, see Fig. 5(a)). When the light is switched off, the tubes cool down and the process is reversed (Fig. 5(b)). Stills from real-time videos [Figs. 5(c) and 5(d) and full video] record the movement of a plasmonic tube under illumination. The dashed white line indicates the starting position of the tube and the red line indicates the position of the tube 1.5 s later (speed 22 lm=s), when the intensity of light is increased to 1 W cm 2. The whole process is completely reversible (see video at Fig. 5). Since the photonic properties of the tubes depend on their morphology, this allows direct control on the sub-lm scale. Such light-sensitive dynamic control of 3D metamaterials not only allows active tuning of the optical properties of the metamaterials but also opens up opportunities for many interesting applications for dynamic rolling/unrolling such as in nano-actuators. The modification of the optical resonances allows mechano-optical feedback so that by irradiating with a specific colour of light which matches a resonant plasmonic absorption (which then selectively heats the Au nanopore film), a particular mechanical state is achieved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000813_amm.760.515-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000813_amm.760.515-Figure8-1.png",
        "caption": "Fig. 8. Star (S 30 UM, S 50 UM) and Oval (O 30 UM, O 50 UM) aplied forcess",
        "texts": [
            " The plate (P 50 UM) and clicket (C 30 UM) were generated with the layers parallel to the machine table, thus the forces will be applied according to the Fig. 6. Fig. 6. Clicket (C 30 UM) aplied forcess FII FP Fig. 2. Clicket (C 30 UM) SLM manufactured part Fig. 5. Plate (P 50 UM) SLM manufactured part Manufacturing Systems and Fig. 7., while the star (S 30 UM, S 50 UM) and oval (O 30 UM, O 50 UM) are built with the layers on a 45 degree angle in reference to the machine table, thus the forces are applied as in Fig. 8. The parts are cutted because the inner hardness is studied. The cuts are made according to the Fig. 6, 7 and 8, with water cooled abrasive disc thus that the internal structure will not be changed. The cutted parts are prepared for the micro hardness test by grinding them with sandpaper and gradually changing the granulation (1500, 2000 and 2500). This grinding is made in order to obtain a very smooth surface and to remove any marks left from cutting and machining. These marks can influence the micro hardness readings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000707_1.3650625-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000707_1.3650625-Figure3-1.png",
        "caption": "Fig. 3 Tooth profiles and data of test gears",
        "texts": [
            " Starting from the gear shaft, the oscilloscope was triggered through a reduction gearing with a ratio of 1 :2 by a synchronizing impulse. ThuB the thermoelectric output from a single-tooth combination was shown repeatedly on the scope and could be examined in detail. Test box lubrication was by means of a jet of a mild additive gear oil (g-A-121) at 2.0 liters per min with a constant temperature of 90 C. Tests with the synthetic oil were conducted with a polyether lubricant (s-106). The temperature-viscosity characteristic of the two test oils is shown in Table 2. The tooth profiles A, C, and L according to Fig. 3 were used. As is shown, only the profile C concerns a tooth profile with equal Journal of Basic Engineering s e p t e m b e r 1 9 6 5 / 6 4 3 Downloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/ on 02/08/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use addendum teeth and with a normal relative sliding velocity vg/v = 0.44. The tooth designs of the profiles A and L have long and short addendum teeth and a high sliding velocity v0/v of 0.67 and 0.764, respectively",
            " If the temperature factor A is plotted versus a term which takes into account the velocity conditions on the highest point of single tooth loading only [15] where A = temperature factor, characteristic for influence of tooth profile, surface roughness, lubricant and gear material T\u201e = (75 + 4.5y'/ !) (deg C) = blank temperature of pinion for idling gears, oil-injection temperature 90 C Note that the curves representing the test results for To m\u00bbx and To min in Figs. 8, 11, 13, 14, and 15 are drawn in accordance with the foregoing relationship. Influence of Relative Sliding Velocity va/v The gear geometry is explained in Fig. 3, and test results in Fig. 9. As the roughness of the gear was approximately constant at all measurements and since all tooth profiles had the same operating pressure angle of ab = 22.5 deg (pressure angle of cutting tool a0 = 20 deg) and the same face width of 6 = 6 mm, the influence of the sliding velocity on the surface temperature could PI \u2014 p-i/i + !>2,/! P i h + (p 2 / i ) ' / 2 no good relationship between the test results appears. This suggests that the surface temperature is also controlled by other geometric influences"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000599_j.engfailanal.2012.02.008-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000599_j.engfailanal.2012.02.008-Figure9-1.png",
        "caption": "Fig. 9. Detail of field of equivalent stresses on pedestal traverse.",
        "texts": [],
        "surrounding_texts": [
            "Computational model of traverse (due to symmetry only one half of beam and middle part with arm of toothing was considered) is given in Fig. 7. Computation of traverse was accomplished for all load cases given above. In Fig. 8 is given field of equivalent stresses on traverse of casting pedestal for the most danger state of loading. In Figs. 9 and 10 are shown details of equivalent stresses on upper side of traverse beam."
        ]
    },
    {
        "image_filename": "designv11_13_0000157_socpar.2014.7007976-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000157_socpar.2014.7007976-Figure5-1.png",
        "caption": "Fig. 5. Photo of the experimental test rig from CWRU, composed of a 2 hp motor (left), a torque transducer/encoder (center), load (right. The test bearings support the motor shaft [23].",
        "texts": [
            " The EMDBB based method combines the EMD and the bispectrum of the defective bearing non-stationary vibration signal. On the other hand, EMDBB can detect sets of frequency components that are phase-coupled. IV. EXPERIMENTAL RESULTS In this section, the BSEMD approach described above is applied to fault diagnosis of ORF BDs. The vibration data of roller bearings analyzed in this paper comes from Case Western Reserve University (CWRU) bearing data center. The detailed description about the test rig can be found in [23]. As shown in Fig. 5, the test rig consists of a 2hp, three-phase induction motor (left), a torque transducer (middle), and a dynamometer-load (right). The transducer is used to collect speed and horsepower data. The load is controlled so that the desired torque load levels could be achieved. The test bearing supports the motor shaft at the drive end. Single point faults with fault diameters of 0.1778 mm, 0.3556 mm, and 0.5334 mm respectively, were introduced into the test bearing using electro-discharge machining. Vibration data are collected using an acquisition system at a sampling frequency of 12 kHz for different bearing conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002435_iros.2015.7353998-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002435_iros.2015.7353998-Figure2-1.png",
        "caption": "Fig. 2. Bicycle of a bevel tip needle",
        "texts": [
            " (5) Solving (5) for a constant forward velocity input u1 results in the position of point P following a circle in Y \u2212Z plane with radius `(tan\u03b2) \u22121. A bevel tip needle inserted into soft tissue is usually driven with two inputs, namely longitudinal insertion and axial rotation. During insertion, as a result of tissue reaction forces, the needle bends in 3-D space. In [10], it is shown that the needle tip posture i.e., position and orientation, resembles the posture of a bicycle moving on circular planar path with the insertion velocity acting as the riding speed. Figure 2 illustrates this model of a bevel tip needle with the associated bicycle wheels. In this figure, frames {B} and {C} represent the moving body frames attached to the wheels. The parameters `1 and `2 denote the distance between the two wheels and the distance between the back wheel and the needle tip, respectively. In this model, the insertion velocity is equal to wheel rolling velocity vz in the body frame {B}. Due to the planar motion of the bicycle, the velocity of frame {B} along its x axis is zero"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002322_j.triboint.2015.10.007-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002322_j.triboint.2015.10.007-Figure9-1.png",
        "caption": "Fig. 9. Installation of the water-sealing system in the holder.",
        "texts": [
            " When a measurement limitation of 1000 Nmm, water leakage of 4250 ml, or abnormal vibration of the shaft seal was observed, the shaft seal was deemed to have failed, and the experiment was stopped. The frictional limitationwas derived from a measurement limit of the torque meter and the leakage limitation was determined by the volume of the pan. 2.4.1. Operating conditions of the water-sealing system The objective of this experiment is to determine the operating conditions of the water-sealing system, especially the appropriate pressure of the PEG lubricant between two hydrated seal lips. The water-sealing system was installed in the experimental apparatus, as shown in Fig. 9. The hydrated seal lips were pressed into a holder (outside diameter: 38 mm, width: 4 mm). The frictional torque and water leakage may be influenced by the amount of pressure exerted by the PEG lubricant between the two seal lips. To examine this influence, two experimental conditions were investigated (Table 1). One was a pressurized condition, in which value (b) was open and the pressurized lubricant was supplied between the two seal kips (Fig. 7). The other was a nonpressurized condition, in which valve (b) was closed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001494_01691864.2018.1493397-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001494_01691864.2018.1493397-Figure5-1.png",
        "caption": "Figure 5. Modeling of an actuator unit.",
        "texts": [
            " M(z) = [ Mw(z) Mr(z) ] = \u23a1 \u23a2\u23a2\u23a3 M11 0 M13 M14 0 M22 M23 M24 M31 M32 M33 0 M41 M42 0 M44 \u23a4 \u23a5\u23a5\u23a6 , (2) M11 = M22 = mP + mC + mB, M33 = IP + IC + mC(xC2 + yC2) + mBr2, M44 = mB, M14 = M41 = mB cos \u03b8 , M13 = M31 = \u2212mC(xC sin \u03b8 + yC cos \u03b8) \u2212 mBr sin \u03b8 , M23 = M32 = mC(xC cos \u03b8 \u2212 yC sin \u03b8) + mBr cos \u03b8 , M24 = M42 = mB sin \u03b8 , H(z, z\u0307) = [ Hw(z, z\u0307) Hr(z, z\u0307) ] = \u23a1 \u23a2\u23a2\u23a3 0 0 H13 H14 0 0 H23 H24 0 0 H33 H34 0 0 H43 0 \u23a4 \u23a5\u23a5\u23a6 , (3) H13 = \u2212mC\u03b8\u0307 (xC cos \u03b8 \u2212 yC sin \u03b8) \u2212 mBr\u0307 sin \u03b8 \u2212 mBr\u03b8\u0307 cos \u03b8 , H14 = \u2212mB\u03b8\u0307 sin \u03b8 , H23 = \u2212mC\u03b8\u0307 (xC sin \u03b8 + yC cos \u03b8) + mBr\u0307 cos \u03b8 \u2212 mBr\u03b8\u0307 sin \u03b8 , H24 = mB\u03b8\u0307 cos \u03b8 , H33 = mBrr\u0307, H34 = mBr\u03b8\u0307 , H43 = \u2212mBr\u03b8\u0307 , g(z) = [ gw(z) gr(z) ] = \u23a1 \u23a2\u23a2\u23a3 0 g2 g3 g4 \u23a4 \u23a5\u23a5\u23a6 , (4) g2 = (mP + mC + mB)g, g3 = ( mC(xC cos \u03b8 \u2212 yC sin \u03b8) + mBr cos \u03b8 ) g, g4 = mBg sin \u03b8 . Next, the dynamics of the wire-driven part is described. Considering \u03b1ctr = (\u03b1ctr 1 ,\u03b1ctr 2 )T as the control input of the wire tension, the following actuator dynamics is obtainable in the wire length coordinates (see Figure 5): \u03b1ctr = \u03b1 \u2212 Aq\u0308 \u2212 Bq\u0307, (5) where A = diag(a1, a2) and B = diag(b1, b2) denote the inertial and viscosity matrices of the actuators, and vector \u03b1 = (\u03b11,\u03b12) T is the resultant wire tension that acts on the target plate. These two tensions, \u03b1ctr and \u03b1, are different because the resultant \u03b1 is affected by the actuators\u2019 dynamics in Equation (5). For vectors \u03b1ctr and \u03b1, the direction pulled by the wire is defined as positive. Therefore, the positive direction of the wire length qi is the opposite of that of wire tension \u03b1i because the positive tensile force reduces the wire\u2019s length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003778_1.4042636-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003778_1.4042636-Figure3-1.png",
        "caption": "Fig. 3 Spherical joint clearance with contact",
        "texts": [
            " When the clearances are not considered, the center of the sphere is coincident with the center of the ball sleeve and the degree-of-freedom is 3. It is assumed that there is a clearance at the spherical joint that between the third driving limb and the moving platform. The center of the sphere is not coincident with the center of the ball sleeve due to the existence of clearance, which causes the sphere can freely move in the ball sleeve and the spherical joint has six degrees-of-freedom. Constraint of three contact forces replaced constraint of three original motions. As shown in Fig. 3, S3 is the center of the ball sleeve, b3 is the center of the sphere, RS3 is the radius of the ball sleeve, Rb3 is the radius of the sphere, g is the contact plane of the two component of the spherical joint with clearance, and Ann and Att represent normal vector and tangent vector, respectively. In this study, \u201c*\u201d is used to distinguish the parameters with clearance. The coordinate of the center point of moving platform in fixed coordinate system fAg in ideal status can be expressed as APBO \u00bc \u00bdAPBOx APBOy APBOz T (1) Ideally, the rotation matrix that the fBg relative to fAg is written as A BR 5 cb 0 sb sasb ca sacb casb sa cacb 2 4 3 5 (2) 041010-2 / Vol",
            "org/about-asme/terms-of-use The coordinate of the center point of the moving platform when considering the spherical joint clearance in fAg can be expressed as AP BO \u00bc \u00bdAP BOx AP BOy AP BOz T (3) Rotation matrix that fB g relative to fAg can be expressed as A BR \u00bc cb 0 sb sa sb ca sa cb ca sb sa ca cb 2 4 3 5 (4) Eccentricity vector of the clearance is obtained as Ae\u00bc APb3 AP S3\u00bc APBO \u00fe A B RBPb3 AP BO A B R BPS3 (5) Eccentricity scalar of the clearance is expressed as e \u00bc ffiffiffiffiffiffiffiffiffiffiffi AeTAe p \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Ae\u00f01; 1\u00de2\u00fe Ae\u00f02; 1\u00de2\u00fe Ae\u00f03; 1\u00de2 q (6) Normal unit vector of the contact plane g is written as Ann \u00bc Ae e (7) 2.2 Relative Normal Velocity and Relative Tangential Velocity Between Joint Elements. Whether the sphere and the ball sleeve are in contact or not can be expressed as d \u00bc e c (8) d < 0; no contact d \u00bc 0; start contact or start separation d > 0; contact and elastic deformation 8>< >: Where c \u00bc RS3 Rb3 is clearance value. As shown in Fig. 3, C3 on the ball sleeve and D3 on the sphere are contact points when the sphere and the ball sleeve are in contact, so the position coordinate of this two points in {A} can be written as APc3\u00bc AP S3 \u00fe Rb3 Ann (9) APD3\u00bc APb3 \u00fe Rb3 Ann (10) Then, the velocity of contact points can be expressed as AVC3\u00bc AV S3 \u00fe RS3 A _nn (11) AVD3\u00bc AVb3 \u00fe Rb3 A _nn (12) Where AV S3\u00bc Ax B Ar S3\u00fe AV BO (13) AVb3\u00bc AVS3\u00bc AxB ArS3\u00fe AVBO (14) The relative normal velocity AVn and the relative tangential velocity AVt of joint elements with clearance can be obtained by projecting the relative velocity of contact points onto contact plane and normal plane, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001900_icra.2015.7140018-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001900_icra.2015.7140018-Figure1-1.png",
        "caption": "Fig. 1. (a) Assembled FWMAV. (b) Solidworks model of the FWMAV and the wing stopper. (c) Single wing test setup. (d) The diagram of flapping wing actuation system consists of motor with voltage input u, resistance Ra, inductance La, back EMF e, motor moment of inertia Jm, damping Bm, and angle \u03c6m, gear with gear ratio Ng , gear moment of inertia Jg , torsion spring with spring constant Ks, and wing with stroke angle \u03c6, and rotation angle \u03c8.",
        "texts": [
            " The experiments showed that ARC was able to achieve excellent tracking of various wing trajectories (with varying amplitude, bias, frequency and split-cycles) and excellent uncertainties rejection performance. When the system parameters were modified by swapping the original identified wing to a wing with different parameters that are unknown to the controller, the ARC showed no performance degradations, compared to 978-1-4799-6923-4/15/$31.00 \u00a92015 IEEE 5852 a benchmarking PID controller. Compared with the openloop method, ARC showed improved force generation results. The experimental setup for single wing testing platform and the assembled FWMAV is shown in Fig. 1(a)(b). The flapping wings are directly driven by two 2.5 gram, 6mm brushless DC motors (FAULHABER, Clearwater, Florida USA) coupled with torsional springs for kinetic energy restoring. Using a gear transmission, the motor was designed to generate an overall reciprocal motion of the wing. The frame structure, wing stopper, wing, and spring holders were prototyped by 3D printing using a multipurpose transparent resin. A portion of the gear on the load shaft was removed to reduce the weight and moment of inertia",
            " The aerodynamic drag Bs2|\u03c6\u0307|\u03c6\u0307 is estimated based on a quasi-steady aerodynamic model using blade element theory (BET) [1], and CD is the mean drag coefficient averaged over one wing stroke estimated by [1] CD = 1.92\u2212 1.55cos(2.04\u03b1\u2212 9.82), where the angle of attack \u03b1 is assumed to be fixed at 45 degrees. The input gain is Ku = Ng Ka Ra . Note that the quasi-steady model is used here due to the lack of simple closed form for unsteady aerodynamic models [1]. The actuation system diagram for one wing is shown in Fig. 1(d). For controller design and control performance evaluation, it is necessary to precisely quantify the high sensitivity of the force and torque generation to wing kinematics. The strokeaveraged forces and torques under consideration are Fz , Fx, roll torque Tx, pitch torque Ty and yaw torque Tz defined similar to [16] as shown in Fig. 2. With a fixed angle of attack \u03b1, the flapping kinematics of each wing are uniquely defined through its stroke angle, which is assumed to be generated by \u03c6i = Aicos ( 2\u03c0ft 2\u03c3i + \u03c8i ) + \u03c60i, if 0 \u2264 t < \u03c3i f Aicos ( 2\u03c0ft\u22122\u03c0 2(1\u2212\u03c3i) + \u03c8i ) + \u03c60i, if \u03c3i f \u2264 t < 1 f (2) where i represents the right (i = r) and left wing (i = l), Ai is the flapping amplitude, \u03c8i is the phase angle, \u03c60i is the bias angle, and \u03c3i is the split cycle parameter",
            " The resulting error shows a slight offset from 1deg to 1.5deg, but is still very small. As a comparison, this 1V input results in a larger angle offset around Ku/Ks \u2248 9deg in openloop experiments. Force measurement was performed using a six component force/torque transducer (Nano17, ATI Ind. Automation). Due to limited resolution of Nano17 (0.3g resolution on the force and 1/64Nmm resolution on the torque measurement), a rigid 150mm beam setup was used to amplify the lift measurement as shown in Fig. 1(c). The improved resolution was about 0.0106g. The force sensor and beam setup was calibrated with precision weights of 0.1g, 0.5g, 5g and 20g and verified the resolution of at least 0.03g. When calculating the timeaveraged force, sufficient large number of wing-beat cycles at steady state ware used to guarantee the reliability of the results. To test the performance of the lift generation from ARC controlled kinematics, lifts at various amplitudes and frequencies were measured and the average lifts obtained are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002867_arso.2016.7736297-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002867_arso.2016.7736297-Figure6-1.png",
        "caption": "Figure 6. Finite Element Analysis of the equivalent model",
        "texts": [],
        "surrounding_texts": [
            "The offline programming software developed for the blade path planning was made by redeveloped EFORT\u2019s software EF-Robot Studio on the basis of studies of the robot trajectory planning, kinematics and off-line programming and other related knowledge. Then the blade was processed using the robotic files generated by offline programming software, and the test results were analyzed. The virtual work site was constructed in software according to the real environment and the structure of robot polishing system, in order to improve work efficiency and find problems in the program. The actual relative position of each device in the software was determined by the robot's point calibration in real environment. According to the planning methods and principles represented before, the upper surface of the blade polishing trajectory was planned in the EF-Robot Studio. The results were shown in the figure 7 as below. In accordance with the previous set, took \u0394=0.02mm, \u03b4 =5mm, and the blade could be planned for the 9 horizontal trajectory, a total of 850 robot processing points. The flexible robotic polishing system was set up as shown in Figure 9, the parameters of various devices used in this experiment were in the Table II: Took the robot files generated by EF-Robot Studio into the robot operation system, and verified the results of the experiment. After the blade was polished, the effect was as shown in Fig.10 following. Comparing different polishing schemes, the roughness, waviness and other indicators of proposed method have greatly improved relative to conventional manual teaching method. Roughness Ra in 0.8, compared with the traditional method of Ra1.6 has much significant improvement, and the preparation time of the robot program could be reduced by 80%, greatly improving the processing efficiency."
        ]
    },
    {
        "image_filename": "designv11_13_0000883_1350650115593962-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000883_1350650115593962-Figure2-1.png",
        "caption": "Figure 2. Contact between a rough sphere and a smooth plane.",
        "texts": [
            " The solution of equations (2) to (5) with the mentioned technique developed by Venner and Lubrecht can determine the local gap and pressure distribution by elastic material description only. Contact calculation with elastic material behavior To analyze a local load between the contacting bodies depending on the surface roughness, a simple geometrical problem with a well-known analytical solution for a smooth case is studied in this paper. A rough sphere with a defined global normal force is pressed against a smooth surface (see Figure 2). Both bodies are considered as deformable. The used calculation method is based on the elastic half space, which calculates the contact between the elastic sphere and a rigid flat in the current case. Therefore, before the simulation the reduced radius and the reduced modulus were determined according to the Hertzian theory.1 In this special case of a sphere contacting a flat surface, the reduced radius according to Johnson\u2019s book is equal to the radius of the sphere.1 For this numerical contact analysis both the linearly elastic as well as the linear elasticperfectly plastic material description is used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000816_s11431-015-5974-1-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000816_s11431-015-5974-1-Figure6-1.png",
        "caption": "Figure 6 Time bias and network path delay.",
        "texts": [
            " Assuming that server time is accurate, the bias between server and client is denoted by  , the path delay from client to server is 1 , the path delay from server to client is 2 , and the total path delay is  . Three equations are formulated under these considerations: 2 1 1,T T     (5) 4 3 2 ,T T     (6) 1 2 .    (7) If the path delays from client to server and from server to client are presumably equal, that is, 1 2 / 2    , then the three equations above can be rewritten as follows: 2 1 / 2,T T     (8)    4 3 / 2 .T T (9) Then, [(T 2 T1) (T3 T 4)] / 2,     (10)     ( 4 1) ( 3 2).T T T T (11) Thus, the time difference between client and server can be represented by Figure 6. As mentioned previously, although BDS can facilitate highprecision time synchronization and solve the security problem posed by the GPS timing source, problems or deficiencies remain when BDS is used in the time synchronization of communication networks. When BDS is used for large-scale distributed system time synchronization, the system must be divided into multiple network segments, and a time service node must be set for each segment. A BDS receiver and clock interface device must also be applied to that node to synchronize different network clocks"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000413_s12206-011-1201-6-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000413_s12206-011-1201-6-Figure10-1.png",
        "caption": "Fig. 10. Finite element model of the bearing-rotor system.",
        "texts": [
            " 8 shows the first two modal shapes of the air blower. According to TMM, the first critical speed is 2167 rpm, and the second critical speed is 4819 rpm. Fig. 9 shows the unbalance response amplitude of the center of air blower\u2019s impeller. It can be seen that at the rotating speed of 945 rpm, the first unbalanced response amplitude at the center of the air blower is 40.96 10 m\u2212\u00d7 , and at the rotating speed of 1790 rpm, the second unbalance response amplitude is 41.17 10 m\u2212\u00d7 . Ansys11.0 is used for modal analysis to the same air blower rotor. Fig. 10 shows the FEM model of the air blower rotor. Solid 45 is used, which has 8 nodes and 6 freedoms. Combine14 is used to simulate the tapered roller bearing and axial freedom of the system is limited. Since the geometry is complex, the element shape in mesh process we used is Quad. Standard method has been chosen to solve the problem. In this finite element analysis, tapered roller bearings are simulated by Combine14, and axial freedom of the system is limited. The governing equations of the whole FEM model is [ ]{ } [ ]{ } [ ]{ } { }M x K x C x F+ + =&& & (17) where [ ]M denotes mass matrix\uff0c [ ]K denotes stiffness matrix\uff0c [ ]C denotes damping matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002173_detc2014-35099-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002173_detc2014-35099-Figure3-1.png",
        "caption": "Figure 3 Layered view of the dome (a) CAD profile (b) CAD model",
        "texts": [
            " Figures 3 & 4 shows a cusp height variation for 8 top layers with 0.254mm layer thickness leading to an undesirable volumetric error with respect to higher build angles. Figure 5 shows a generalized approach that uses 4 layer thicknesses to demonstrate more effective cusp height variation along with different build angles with the help of mathematical equation to avoid maximum volumetric error conditions or resulting poor surface issues. First approach has been demonstrated with the example part presented in Fig. 3 of this paper. Figure 3 shows the layered view of the dome model and corresponding cusp heights and surface complexity angles are presented in Figs. 4. It clearly shows the interface between the variation of surface slope and the cusp height. In Fig. 3, there are two different staircase regions near steep slope for top two layers. The area enclosed for these top two layers i.e. layer 8 and layer 7 has maximum difference (refer Figs. 3 & 4). From Figs. 3 & 4 it is clear that in this case there will be significant volume loss. However cusp height is almost the same in this situation. Therefore, in this situation cusp height based adaptive slicing algorithm will give almost same layer thickness. Therefore, a comprehensive solution to the surface finish improvement will not be found when cusp height concept is involved in the optimization of steep surface conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002594_1077546314558133-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002594_1077546314558133-Figure1-1.png",
        "caption": "Figure 1. A rotor-bearing system: (a) A model for the symmetrical flexible rotor, (b) Forces applied on the journal and the disk.",
        "texts": [
            " This paper is structured as follows: Section 2 introduces the nondimensional form of the nonlinear model of the flexible rotor bearing system. The next section, Section 3, gives a brief introduction to the method of numerical continuation. Section 4 presents the main results and a discussion of the findings presented in this paper. The paper is concluded in Section 5. Considering a flexible rotor supported by two identical and aligned journal bearings rotating at an angular frequency !. The rotor is composed of a flexible shaft of negligible mass and stiffness 2K and a central perfectly balanced rigid disk of mass 2M, Figure 1(a). Both journal bearings are assumed to have the same motion. The position of the journal centre Oj with respect to the bearing centre Ob is defined by the eccentricity e \u00bc c\" and the attitude angle as shown in Figure 1(b), where c is the bearing clearance and \" is the eccentricity ratio. The position of the disk centre Od with respect to the bearing centre Ob is defined by the Cartesian coordinates \u00f0x, y\u00de. The forces applied on the disk are the disk at East Tennessee State University on May 30, 2015jvc.sagepub.comDownloaded from weight W and the restoring force fr ! applied by the flexible shaft. Applying Newton\u2019s second law yields the following two equations of motion of the disk: M \u20acx\u00fe K\u00f0x c\" sin \u00de \u00bc 0 M \u20acy\u00fe K\u00f0 y\u00fe c\" cos \u00de \u00bc 0 \u00f01\u00de The fluid film forces applied on the journal are represented by the two components f\" "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000655_c2sm26295a-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000655_c2sm26295a-Figure2-1.png",
        "caption": "Fig. 2 Electrowetting experimental set-up: voltage is applied between the fiber tin core and droplet. Contact angle is measured within a 200 mmwindow around the contact line in order to obtain a good accuracy.",
        "texts": [
            " Electrowetting experiments were set up on a drop shape analysis system (DSA100, Kruss GmbH, Germany) and performed with 1 0.1 mL deionized water drops. Frequency of the applied square voltage was xed to 1 kHz for all experiments (lower frequencies favor electrolysis and large frequencies decrease efficiency of the EWOD effect23). All measurements were carried out 5 times at different positions along the ber. We rst performed static EWOD characterizations in order to capture variations of contact angle q(Vrms). An experimental setup is presented in Fig. 2: voltage was applied to the ber tin core and the water drop was connected to the ground through a thin gold wire (25 mm in diameter) so that it did not inuence the drop shape. The applied voltage was continuously varied from 0 to 160 Vrms and contact angles were measured within a 200 mm window in the vicinity of the contact line. Measurement of contact angles of clam-shell drops hanging on horizontal bers is not straightforward compared to classic studies on smooth or textured surfaces.24,25 Because of the complex drop shape, the interface prole continuously varies from amicroscopic contact angle to the macroscopic drop prole (see Fig",
            " In that case the applied 1 kHz square voltage was successively switched on and Soft Matter, 2013, 9, 492\u2013497 | 493 Pu bl is he d on 2 6 O ct ob er 2 01 2. D ow nl oa de d by H un te r C ol le ge o n 12 /0 6/ 20 14 2 1: 22 :3 0. off at a frequency of 0.5 Hz and cyclic variations of contact angle were recorded (5 measurements per second). Maximal hanging drop volumes Uc were obtained by pumping up a pendant drop at a constant applied voltage until falling-off. For this experiment the drop was initially in the conguration depicted in Fig. 2 and a thin capillary was positioned vertically right above the ber. Tiny volumes of liquid dv z 0.3 mL were added by creating micro-drops outside the capillary. Once the micro-drop touched the larger one, it spontaneously coalesced because of Laplace pressure difference, which enables noncontact pumping of the droplet. Right before falling-off, Uc was estimated using rotational symmetry approximation (\u2018pendant drop\u2019 method, DSA3 soware, Kruss GmbH, Germany). We assume that this method leads to about 5% overestimation ofUc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003961_s42417-019-00132-1-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003961_s42417-019-00132-1-Figure2-1.png",
        "caption": "Fig. 2 Ball rotation vector in moving coordinates",
        "texts": [
            " This technique of deduction of the mathematical model in a particular case consists mostly in conducting of geometric relations among motion components and, finally, in construction a detailed form of the Appell\u2013Gibbs function. Subsequently, by means of differentiation with respect to adequate response components, the governing differential system can be carried out. The Appell\u2013Gibbs function (often referred to as an energy acceleration function) is defined as a function of six component characterizing motion of one stiff body in 3D: where it has been denoted (see also Fig.\u00a02): M is the mass of the ball, J is the central inertia moment of the ball with respect to point G, is the angular velocity vector of the ball with respect to its center G, G = ( x, y, z) is the displacement of the ball center with respect to absolute origin O, C is the contact point of the ball and cavity, A is the moving origin related to the cavity in its bottom point, = (x, y, z) is the Cartesian coordinates with origin in the point O. It should be noted that the general denomination of the ball characteristics M,\u00a0J later allows to consider particular cases whatever type of spherical body with central symmetry, the mass of which is either concentrated in the center (J = 0) , uniformly distributed within the body (J = 2\u22155Mr2) , or evenly dispersed over the outer envelope of the ball (J = 2\u22153Mr2)",
            " Concerning the latter one, complicated energy dissipating processes act in the contact of the ball with the cavity. Nevertheless, supposing that no slippage arises in the contact, the dissipation process can be approximated as proportional to relevant components of the angular velocity vector and the quality of the cavity/ ball contact. Considering the obvious setting, the respective material coefficients characterizing the rolling movement of the ball can be considered constant regardless the direction in the tangential plane to the cavity in the point C, see Fig.\u00a02. The coefficient determining the rotation resistance around the normal vector in the contact point C (spin rotation component) is different as a rule, because the spin rotation is related rather to a dry friction. Nevertheless, influence of this damping force is even smaller than those acting in (17) u\u0307Cx = \ud835\udf14y(uCz \u2212 R) \u2212 \ud835\udf14zuCy u\u0307Cy = \ud835\udf14zuCx \u2212 \ud835\udf14x(uCz \u2212 R) u\u0307Cz = \ud835\udf14xuCy \u2212 \ud835\udf14yuCx. (18) ?\u0307?x = P1\u2215Js, ?\u0307?y = P2\u2215Js, ?\u0307?z = P3\u2215Js. 1 3 tangential directions and, therefore, such an approximation is acceptable. Consequently, the resistance moment vector can be expressed in moving coordinates p,\u00a0q,\u00a0n, see Fig.\u00a02, as follows: Components of the above vector can be written as follows: where , are coefficients of \u201cviscous resistance\u201d of rolling and spinning. Their meaning is: the moment for a unity rotation per second, i.e., [Nms/rad]. Turning of the vector G = [DGx,DGy,DGz] T expressed in (xyz) coordinates into the vector can be written as follows: The transformation matrix TC reads the following: where 2 = x2 C + y2 C . The matrix C is orthogonal and, therefore, the inverse transformation is easy to obtain: \u22121 C = T C ",
            " Of course, in such a case, some results have rather the theoretical meaning only. Nevertheless, such results are still important for completeness and consistency of the response portrait of the system as a whole. Before we start to discuss individual series of ball trajectory types, some initial remarks are useful to be done. It regards primarily the basic Separating Horizontal Circle (SHC), which enables to relate all the others to that as the reference point. In general, it is horizontal circular trajectory which starts in a certain point C, see Fig.\u00a02. The ball is elevated into this point and thrown horizontally along a parallel line of the cavity under assumption that no damping acts in the contact ball cavity. The relevant impulse is realized by the initial vector = ( p0, 0, 0) passing the ball center G, which is parallel with the tangent to the meridian in the point C. Due to constraints [Eq.\u00a0(17)], it represents a horizontal movement in the starting moment. The initial value p0 = ps should be so determined that the horizontal direction is kept"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001250_j.matpr.2017.07.234-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001250_j.matpr.2017.07.234-Figure8-1.png",
        "caption": "Figure 8 The integrated substrate at the centre of component",
        "texts": [
            "15603 Hence the substrate size will be 2 75, 2 235, 5 Sajan Kapil / Materials Today: Proceedings 4 (2017) 8837\u20138847 8843 Now in this orientation the position of substrate has to find out such that the buildability of the overhanging feature could be possible. It has been found that the best fit without sharp overhanging feature is in the zone 5, is half way between top and the bottom planes. The next step is now to find out the maximum intersecting volume position in +Z (build direction) direction. As discussed earlier that the zone for inserting the substrate is 5 hence it will be convenient to transfer the work plane at the centre of the component. In Figure 8 the substrate is placed at the centre which is now the position 0 and the zone for finding the maximum intersection will be 5. The maximum engagement of the substrate with the component has been found by sliding it in direction. Here, the work plane of the substrate has been shifted by 1 mm in direction. Then, volume occupied by the substrate of the component has been found by Boolean operation. Table 2 shows the volume of the component covered by substrate at different level. Here (+ ) layer position describes above the centre level and (- ) layer position describes below the centre level"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001774_978-3-642-39739-4_15-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001774_978-3-642-39739-4_15-Figure6-1.png",
        "caption": "Fig. 6 FCM for selecting an object of shooting the ball: a structure and b inference scheme",
        "texts": [
            " Just on this level rules for calculating suitability are placed. Finally, the third level is created by the set of possible actions, e.g. shooting, passing or dribbling the ball. The action, whose node achieves the highest activation value, will be chosen. Another design for an action selection FCM is described in [24], which is also constructed for purposes of robotic soccer. In this case it serves for choosing the object of shooting the ball. Either it is a direct shot to the opponent\u2019s goal (G) or passing to a teammate (TM), see Fig. 6. The aim of the object selection is minimization of the risk that the ball will be caught by the opponent. The structure of such an FCM is in Fig. 6a. Each teammate has its operational range, i.e. area, which can be reached during a time step and is described by membership functions in Fig. 6b. Also the goal G (although static) can be described in such a manner because it occupies a certain area. The only difference is that because of the dynamic nature of teammates (they move during the play) their membership functions must be newly generated each time step (the same for the closest opponent O). Also the weights of connections wsi (Fig. 6a) change over time because they are indirectly proportional to the distances between the shooter and objects (the closer the smaller risk to be caught by the opponent). Peaks in the membership functions of the nodes TMi represent angles of view for objects seen from the shooter\u2019s perspective. Values of nodes Aj are also indirectly proportional to the size of the turn angle needed for the shooter. These nodes represent possible angular ranges (intervals) of shooting to individual objects for a given shooter dividing the total range of 360 \u25e6 and calculating the cost of turning for the shooter to a given range. The calculation of resultant kicking angle is calculated in the node S. As membership functions in nodes TMi and G should help attract the ball and the function in node O should repel the ball, this function will be constructed in a manner reverse that of other functions, i.e. it will be minimized at the measured angle of the opponent, see Fig. 6b. The values of nodes Aj will then be multiplied together with weights wsi , yielding total costs of shooting that include turning and distance. These costs will be used to scale functions TMi and finally all membership functions will be united in the node S using a sum or more correctly a t-conorm. The first maximum of such a final membership function determines the kick angle and its closest object will be selected (in Fig. 6b it is the teammate TM1). Although FCMs show many excellent properties but they are of course no nostrum. Beside further high quality approaches also lots of highly specialized methods have been developed and against them means for general purposes cannot compete in their area of use. Therefore, the idea of hybridization plays in some applications an important role. Principally, we can distinguish two ways of hybridization, either FCM separately from other means cooperates with them as an element of a compound system or the structure of FCM absorbs further means"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001977_j.taml.2015.08.003-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001977_j.taml.2015.08.003-Figure1-1.png",
        "caption": "Fig. 1. A three-mass chain (a) and its equivalent two-mass system (b).",
        "texts": [
            " The idea is the following: we consider a three-mass chain, one adjacent pair of the masses is designed to behave as a body with iety of Theoretical and Applied Mechanics. This is an open access article under the CC equivalent negative mass, and the third positive mass interacts with this equivalent negative mass through an elastic spring. Both theoretical analysis and experiment are performed to illustrate the possible motion mode of the diametric drive. We will also examinewhether Hamilton\u2019s principle still holds for a systemwith an effective negative mass concept. Consider a three-mass chain, as shown in Fig. 1(a), both the mass mi and stiffness kj (i = 0, 1, 2 and j = 1, 2) are positive. The mass 2 is connectedwith themass 1, under harmonicmotion, these two masses may be considered as an effective mass connected to the mass 0, as shown in Fig. 1(b). Using Newton\u2019s second law, the equations of motion for the three masses are written respectively as m0x\u03080 = k1(x1 \u2212 x0), (1a) m1x\u03081 = \u2212k1(x1 \u2212 x0) + k2(x2 \u2212 x1), (1b) m2x\u03082 = \u2212k2(x2 \u2212 x1), (1c) where xi(t) is the displacement of the mass i. If we consider a steady harmonic motion e\u2212i\u03c9t and consider m2 as a hidden mass, we eliminate x2 from the equations, this leads to m0x\u03080 = k1(x1 \u2212 x0), (2a) meff x\u03081 = \u2212k1(x1 \u2212 x0), (2b) where meff = m1 + m2/(1 \u2212 \u03c92/\u03c92 2), and \u03c92 = \u221a k2/m2. Eqs. (2a) and (2b) are exactly the same for a two-mass system, i.e., m0 and meff are connected by the spring k1. Therefore we can tune the frequency\u03c9 to letmeff < 0, and examine in turn its interaction with the positive mass m0. For a general two-mass system with one negative mass, Newton\u2019s second law tells that depending on the absolute value of the negativemass, the systemwill oscillate if it is larger than the positive one and accelerate in one direction otherwise. For the system shown in Fig. 1, it has two natural frequencies, satisfying the following equation \u03c9\u03044 \u2212 (\u03c92 1 + \u03c92 2 + \u03b12)\u03c9\u03042 + \u03c92 1\u03c9 2 2\u03b2 = 0, (3) where \u03c91 = \u221a k1/m1, \u03b1 2 = k2/m1 + k1/m0 and \u03b2 = (m0 + m1 + m2)/m0. So the two natural frequencies are given by \u03c9\u03042 1,2 = 1 2  \u03c92 1 + \u03c92 2 + \u03b12 \u00b1  (\u03c92 1 + \u03c92 2 + \u03b12)2 \u2212 4\u03c92 1\u03c9 2 2\u03b2  . (4) At the lower nature frequency \u03c9\u03041, the hidden mass m2 moves out of phase with respect to the mass m1, leading to that the pair of the masses m1 and m2 behaviors as an effective negative mass, i.e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001673_s11837-018-3242-0-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001673_s11837-018-3242-0-Figure5-1.png",
        "caption": "Fig. 5. Magnetic flux density for 20-W BLDC PMM with anisotropic bonded magnets aligned at (a) 0.25 T and (b) 1 T.",
        "texts": [
            " This is expected since the magnetic properties of the aligned anisotropic bonded magnets, in particular the remanence, are higher than for the isotropic magnet, which increases the magnetic interaction between the rotor and stator (cogging torque). The primary impact of torque ripple is audible noise, which is not a major concern for submersible water pumps, where such effects are heavily damped by the operating environment.26 Known design techniques such as slot skew and rotor skew could also be employed to reduce the torque ripple.27,28 Control methods can also be employed.29 Figure 5 compares the magnetic flux density in the 20-W SPM BLDC motor for anisotropic bonded magnets aligned at 0.25 T and 1 T. The small increase in remanence provided by the higher alignment field strength did not greatly increase the magnetic flux density in the rotor or stator yoke. Saturation is approached in a small section of the Khazdozian, Li, Paranthaman, McCall, Kramer, and Nlebedim stator teeth for both magnets, but does not occur in the stator or rotor yoke. Flux leakage in the airgap and rotor yoke may account for the nonuniform distribution of magnetic flux density in both cases, as well as for the low efficiency of this small motor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001992_ecce.2015.7310061-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001992_ecce.2015.7310061-Figure6-1.png",
        "caption": "Fig. 6: a) Simplified rotor structure for evaluation of b) d axis MMFFEM",
        "texts": [
            " Therefore solution of is given by The induced eddy currents are: The eddy current loss in each segmented magnet can then be computed as: The eddy current loss in a magnet segment of an IPM machine as given by the above is proportional to the volume of the magnet and its conductivity. It also varies as square of the flux density of the source field. The second term in the denominator of the loss equation, , represents the effect of reaction field due to the induced eddy current in the magnet. Estimation of Consider the simplified picture of the motor geometry as seen in Fig. 6. The value of is estimated using the scalar magnetic potential [20]. Let the MMF of the stator be aligned to the d axis. The instantaneous d axis MMF component of a three phase winding can then be written as a function of circumferential angle as where is the number of pole pairs and is given by number of series turns/phase of the three phase winding , the maximum phase current in the winding and the winding factor as The average d axis MMF across the magnet width of the stator winding can be calculated as: With the magnet and air-gap reluctances and defined as and , the scalar magnetic potential across the magnet is evaluated as The flux density in the magnet can then be evaluated as The above equation gives a good estimate of "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000748_tie.2013.2276025-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000748_tie.2013.2276025-Figure12-1.png",
        "caption": "Fig. 12. Distribution of temperature T in the billet after 180 s of heating for nnom = 1420 r/min.",
        "texts": [
            " 10 shows the distribution of magnetic flux density B in the system for the nominal revolutions of the billet nnom = 1420 r/min. Only very small areas near the corners of joints between the permanent magnets and the magnetic circuit are oversaturated. Analogously, Fig. 11 shows the distribution of the volumetric Joule losses pJ produced by the currents induced in the rotating billet at the same revolutions. In accordance with theory, its highest values are generated in the surface layers of the billet. Fig. 12 shows the distribution of temperature in the billet after 180 s of heating. It is obvious that the temperatures over its whole cross section do not differ more than by a few degrees of centigrade. This is caused by the very good thermal conductivity of aluminum. Fig. 13 shows the distribution of temperature along the radius of the billet in time. Due to varying remanence of the permanent magnets (ranging from 1.38 to 1.45 T), we also investigated its influence on the rate of heating of the billet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000135_icra.2012.6225025-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000135_icra.2012.6225025-Figure5-1.png",
        "caption": "Fig. 5. Some examples for cases considered in the cross section collision check",
        "texts": [
            " Algorithm 2 SS-CollisionCheck(segi, fj) Compute the distance between segi and Qj , denote the distance as d(segi, Qj) and closest points found on segi and Qj as p, q respectively; if d(segi, Qj) \u2264 wi then if q is not in fj then Compute the minimum distance between each edge ejk of fj and segi, denoted as dmin(segi, e j k) with closest points p and q on segi and ejk respectively. if dmin(segi, e j k) > wi then Return Collision = False. end if end if if p is not on pi\u22121 or pi then Return Collision = True. else if q is on Hi\u22121 (or Hi) then Return Collision = True. end if end if end if Return Collision = False. v2 as (\u03c11, \u03b81) and (\u03c12, \u03b82) respectively. Algorithm 3 partitions all scenarios into five cases based on whether v1 and v2 satisfy the bounds of inequalities (1) and (2) to detect intersections (i.e., collisions). Fig. 5 shows examples for these cases. If a face fj of the object does not intersect Pi or the cross section csi, we need to further check if fj intersects section i by Algorithm 4. In Algorithm 4, we first check if the distance between ciri and Qj , the supporting plane of face j, is greater than the width of section i by calling Procedure 1. If so, Qj has no intersection with the section i, a truncated torus, and no further collision checking is necessary. If Qj intersects the section i, then further collision checking is done by calling subsequently Procedure 2, and if necessary, also Procedure 3",
            ", min(\u03c11, \u03c12) > ri + wi then Compute the distance between circle center ci and lij to obtain point q = (\u03c1q, \u03b8q) on lij . if q is within csi, i.e., \u03c1q and \u03b8q satisfy inequalities (1) and (2) respectively then Return Collision = True. else Return Collision = False. end if Case 4: if both \u03b81 and \u03b82 satisfy (2) then Return Collision = True. Case 5: if line segment lij intersects either ray Li,k (k = 1, 2) of csi (see Fig. 4) then if lij is colinear to either ray then Return Collision = True. end if if At least one of \u03c1kint satisfies (1) then Return Collision = True. (Fig. 5(e)) end if if lij intersects both rays at one point above the upper bound for \u03c1 and one point below the lower bound for \u03c1 then Return Collision = True. (Fig. 5(f)) end if if lij intersects only one ray at a point above the upper bound for \u03c1 and the vertex of lij with the smaller \u03c1 value satisfies (2) then Return Collision = True. (Fig. 5(g)) end if if lij intersects only one ray at a point below the lower bound for \u03c1 and the vertex of lij with the greater \u03c1 value satisfies (2) then Return Collision = True. (Fig. 5(h)) end if Return Collision = False. (to the original point) as q\u2032 and the corresponding distance dmin(pi/i\u22121, Qj). \u2022 If q\u2032 is on fj then return dmin(pi/i\u22121, Qj) as dmin(pi/i\u22121, fj). \u2022 Else, find the minimum distance between each point to every edge of fj and return the shortest distance as dmin(pi/i\u22121, fj). Procedure 3: Compute the minimum distance dmin(segi, ek) between segi and an edge ek, and obtain the pair of closest points: \u2022 Project ek to Pi and denote it as eik. \u2022 If eik intersects ciri at point p, then \u2013 If p is on segi, return the distance from p to ek as dmin(segi, ek), as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002608_1464419316636968-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002608_1464419316636968-Figure7-1.png",
        "caption": "Figure 7. Clearance between wheel flange and rail gauge.",
        "texts": [
            "comDownloaded from Kalker\u2019s linear theory was used to calculate the longitudinal and lateral forces and the vertical moment of the two wheelsets. Then, by an assumption of small roll and yaw angles, linear creep forces and linear creep moments were obtained for the left and right wheels. Finally, the heuristic nonlinear creep model was used to consider the nonlinear longitudinal and lateral forces and also the nonlinear vertical moment.27 The lateral displacement required for the wheelset to make flange contact with the rail is known as the flange clearance between the wheel and the rail (see Figure 7). When the lateral displacement of the wheelset is less than the flange clearance, both wheels of the wheelset are in single point tread contact with the rails. Alternately, when it is greater than the flange at UNSW Library on April 9, 2016pik.sagepub.comDownloaded from clearance, one of the wheels makes a single point flange contact with the rail. When the lateral wheelset displacement becomes equal to the flange clearance, both the tread and the flange of the wheel make contact with the rail and a double point contact condition, involving three different contact patches, occurs (two at the flanging wheel and one at the other one)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000281_1.3657244-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000281_1.3657244-Figure3-1.png",
        "caption": "Fig. 3 Contact roller test machine",
        "texts": [],
        "surrounding_texts": [
            "50 microinch. The Ri circularity was monitored by means of Talysurf traces obtained with a fixed radius bar which corresponded to the radius of interest.\nThe teste were conducted in eight contact roller machines, one of which is shown in Pig. 3. The rollers were pressed together under dead-weight load and rotated at 1800 cpm. The cylindrical element was the drive roller in all tests. The rollers were supported in the housings by Typo 303 and 5303 ball bearings. The rollers and bearings were amply lubricated with Navy Symbol 2190 TEP lubricating oil (Socony Mobil RL285D, L-3077). The oil was continuously filtered and maintained at an inlet temperature of 90 F. The lubricant was introduced by jet in copious quantity to the converging surfaces. At the inception of failure, i.e., breakdown of roller surface, the machine was stopped automatically.\nMaterials The contact rollers were fabricated from three heats of bearing quality AISI 52100 steel. The chemical analysis and production history of the steels are shown in Table 2. The macrostructures of all bars showed no unusual or significantly different metallurgical features. The microstructures and the largest inclusions found are shown in Fig. 4. Material D I W contained more oxidetype inclusions than the other steels. The inclusion content approximated the ASTM Type D, No. 2 thin series rating.\nBlanks for roller fabrication were cut from the bars in integral sets; i.e., the test and drive rollers for Run 1 were cut from adja-\ncent material. Thus a set of rollers represented a local unit of bar stock. In addition, blanks were cut so that rollers of various toroid radii were produced from successive units of bar; i.e., rollers of radii 0.250, 0.383, and 0.500 inch were fabricated from Cuts 1, 2, and 3, respectively. This procedure was repeated throughout the heat.\nThe roller blanks were machined to shape allowing 0.015-inch oversize on diameters. The randomly selected rollers then were heat-treated in an Ipsen, Model RT-25-E, unit with a Type AGA301 atmosphere. The following schedule was used (1) Austenitize at 1500 F for 1 hr; (2) quench in oil at 140 F; and (3) temper for IV2 hr at 350 F. The resulting hardness was 62 to 64 Rc. Representative microstructures of the hardened steels are shown in Fig. 5. The hardened rollers were rough-ground to 0.002- in. oversize and finish-ground on a Norton Type CTW cylindrical grinder, using a Stearling WA 180K7V2 wheel. A profile trace of each roller surface perpendicular to the grinding direction was obtained with the Talysurf, Model 3, instrument. Representative profile traces are shown in Fig. 6.\nResults 194 tests of the scheduled 243 were tested as planned. Trunnion breakage caused most of the aborted runs, while machine failure and personnel error caused five aborts. Trunnion failure resulted from the shoulder undercut which was made to permit easier grinding. The heavier-loaded rollers, and especially the longer cylindrical rollers used with the 0.383-in. toroid, showed\n1 8 2 / M A R C H 1 9 6 2 Transactions of the A S M E\nDownloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jfega4/27236/ on 06/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
            "Journal of Basic Engineering\nM A R C H 1 9 6 2 / 1 8 3\nDownloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jfega4/27236/ on 06/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
            "high propensity for fatigue failure in this location. Preliminary runs of rollers had failed to reveal this subsequent difficulty.\nA typical fatigue failure of a contact surface is shown in Fig. 7. The distribution of failures between the mating surfaces is shown in Table 3. The distribution appeared random except for the 0.383-in. roller group loaded with 2014 lb. The cause for this is unknown.\nMetallurgical sections, cut perpendicular to the roll direction\nand well removed from the area of surface failure, frequently revealed material damage that had not yet progressed to the surface. Examples are shown in Fig. 8. Indications of incipient material transformation similar to that observed in elevated temperature tests, reference [1], also were observed. This was seen with the unaided eye after etching with mild reagents. The transformation was not apparent at 100, 250, or 500 magnifications. X-ray diffraction determinations indicated considerable\n184 / MARCH 1 9 6 2 Transactions of the A S M E\nDownloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jfega4/27236/ on 06/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use"
        ]
    },
    {
        "image_filename": "designv11_13_0001136_s00170-017-0625-2-Figure20-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001136_s00170-017-0625-2-Figure20-1.png",
        "caption": "Fig. 20 Profile of the top molten pool for thin-wall samples. The arrows indicate the grain growth direction in this molten pool",
        "texts": [
            " Figure 19 shows the different droplet transfers in the thin wall and ladder block processing (The arrows show the cooling direction in a molten pool). Because of the molten pool boundary from an adjacent bead, the liquid bridge in the ladder block deposition processing is also asymmetric, which induces the versatile grain growth direction in this position. A cooling direction that points to an adjacent bead being produced and more types of grain morphologies are formed under this basis. The microstructure-evolution phenomenon is particularly clearly visible in thin wall depositions. Figure 20 shows the profiles of the top molten pools of each thin wall sample. (The profiles are delineated from Fig. 7). Because of the effects of pulsed laser, several fusion lines are seen within the one-layer deposition. The fusion lines that are present throughout the entire width direction of section X can be regarded as the boundaries between adjacent layers. (The arrows show the grain-growth direction in this molten pool.) In the beginning of the thin wall deposition, the outer contour of section X presents a continuous curve shape, which indicates that the top molten pool almost covers the whole surface of depositions. Consequently, columnar grains grow on both sides of the deposition gradually (Fig. 20). With increasing deposited layers, the curvature of the top surface decreases gradually. In this study, the shape of the top surface starts to become almost horizontal in S13, and the outlines of the top and side surfaces are completely separated. In this time, the base of the molten pool was almost unchanged, and the shape of the molten pool became stable. Consequently, the microstructure morphology does not vary with the deposition process anymore (Fig. 20). Nevertheless, the shape of the molten pool is actually more stable in the ladder block deposition than in the thin wall deposition, which is due to the effect of the overlap process. On the one hand, it makes the top surface of the ladder blocks smoother than that in the thin wall, which is an advantage condition for a stable deposition process. On the other hand, due to the limitation caused by the side boundary of the adjacent bead, space for the liquid metal flow will decrease. Consequently, there is no variation in the microstructures in the same positions of different ladder block samples"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003587_iet-epa.2018.5161-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003587_iet-epa.2018.5161-Figure11-1.png",
        "caption": "Fig. 11 Installation figure of the magnetic field measuring the coil",
        "texts": [
            " 10, and because it is affected by failures in the A1 and A2 branches, the armature magnetic potential will be asymmetric at the conducting time of the A-phase positive half cycle. Once it rotates over one pair of magnetic poles, the armature will become asymmetric for once. Once the armature rotates one cycle, the armature magnetic potential will be asymmetric eight times. To detect an open-circuit fault in the diode of a rotating rectifier in a brushless exciter, two detection coils are installed on the exciter stator's iron yoke with a spacing of 180\u00b0, as shown in Fig. 11. Once rotation occurs over one pair of poles, the phase will change six times. Fig. 12 refers to the armature magnetic field figure during a single commutation. At the time of Fig. 12a, the A and C phases are conducting. The arrows indicated in this figure refer to the armature magnetic field. The armature rotates by a \u03c0/3 to the position shown in Fig. 12b and then the armature magnetic field also rotates by \u03c0/3. After the armature winding rotates by the \u0394 angle (infinitely small), the armature winding changes from conducting in the A and C phases to conducting in the B and C phases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002737_9781118899076-Figure2.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002737_9781118899076-Figure2.1-1.png",
        "caption": "FIGURE 2.1 The statistical detection limit. Reproduced from Ref. [1] by permission of Elsevier.",
        "texts": [
            " During the last decade, the DL of ISEs has been radically improved, that is, lowered. This has led to the prospect of significantly expanding the use of ISEs in such important fields as clinical or environmental analysis. Electrochemical Processes in Biological Systems, First Edition. Edited by Andrzej Lewenstam and Lo Gorton. \u00a9 2015 John Wiley & Sons, Inc. Published 2015 by John Wiley & Sons, Inc. 23 The criterion of detection of any analytical method is defined as the level of the measured signal (point Q in Fig. 2.1) at which the probability of error equals \u03b3. The error is defined as the incorrect analytical conclusion; in this case, the judgment that the determined substance is present in the sample when in reality it is not and the high signal level are caused by random errors. The criterion of detection is defined as Q=\u03a61\u2212\u03b3\u03c3Q 2 1 where \u03c3Q is the standard deviation of a single measurement of the difference between the sample and the blank signals: \u03c3Q = \u03c3 A\u2212B = \u03c32A + \u03c32B 2 2 Assuming that the standard deviations of the signals for the sample and the blank are equal (\u03c3A = \u03c3B), we obtain Q=\u03a61\u2212\u03b3 2 \u03c3B 2 3 If we assume that the probability of error \u03b3 equals 0.05, Q = 1 645 2 \u03c3B = 2 33 \u03c3B 2 4 24 TRANSMEMBRANE ION FLUXES FOR LOWERING DETECTION LIMITS The DL can be defined in an analogical way (point L in Fig. 2.1). In this case, the wrong analytical conclusion will be that the substance to be determined is not present in the sample when in reality it is present, and the small difference between the signals of the sample and the blank is caused by random errors: L=Q+\u03a61\u2212\u03b3\u03c3L 2 5 Assuming that \u03c3A = \u03c3B, we obtain L=Q+\u03a61\u2212\u03b3 2 \u03c3B = 2 2 \u03a61\u2212\u03b3 \u03c3B 2 6 Assuming, as before, that the error \u03b3 = 0.05, we eventually obtain L= 4 65 \u03c3B 2 7 2.2.1.1 Application of the Statistical DL to ISEs The DL of ISEs is defined in a specific way that differs from the definition used for other analytical techniques"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001914_ilt-09-2013-0101-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001914_ilt-09-2013-0101-Figure1-1.png",
        "caption": "Figure 1 Schematic views of bump-type foil bearings",
        "texts": [
            ", are also being studied (DellaCorte, 2012). GFBs are self-acting, hydrodynamic bearings with a flexible underlying foil structure. The underlying supporting structure consists essentially of an overlapping smooth foil that acts as the bearing surface and a corrugated bump foil that provides a flexible support to the top foil. A hydrodynamic air film builds up in the gap between the top foil and the housing which forms from the clearance and the foil deformation. The layout of a typical bump-type foil bearing is illustrated in Figure 1. Misalignment is one of the most common malfunctions in rotating machinery (Muszynska, 2005). This malfunction may be caused by manufacturing tolerances, asymmetric bearing loads, installation errors and thermal distortions. The presence of corrupted foil strips allows GFBs to accommodate higher degrees of misalignments in bearing assemblies than in rolling element bearings (Howard, 2009). To determine the misalignment limits of GFBs, Gray et al. (1981) tested four GFBs with different foil structures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002407_s1064230715040115-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002407_s1064230715040115-Figure1-1.png",
        "caption": "Fig. 1. Path variables.",
        "texts": [
            " The time scale transformations discussed can be applied to more complicated WR models. Since the path following problem consists in stabilizing motion of a certain target point C on the robot platform along a given path, we may always assume that the coordinates of this point in some coordinate system are known and take them to be state variables. For the coordinate system, it is convenient to take the Frenet frame, which is the moving frame with the origin located at the curve point closest to the target point C (Fig. 1). The abscissa axis of the frame is directed along the tangent line to the curve in the positive direction of the curve. The coordinates of the point C in this frame are evidently (0, d), where d is the dis tance between C and the target curve. Given a parametric description of the target curve, the position of the target point in the plane is completely determined by the two variables: d and the value of the path parameter s at the origin. The orientation of the robot platform can be given by the angle between the velocity vector of point C and the abscissa axis of the Frenet frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003504_taes.2018.2852419-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003504_taes.2018.2852419-Figure1-1.png",
        "caption": "Fig. 1. Missile kinematics",
        "texts": [
            " Section IV elaborates how to determine a unique nominal flight course, which is represented by a nominal heading angle and a nominal heading angular rate, by utilizing the properties of the nominal motion. Section V gives the continuous backsteppingbased finite time guidance law that yields the finite-time tracking of the heading angle and of its rate. In Section VI, a series of simulation results demonstrates the validity of our method. Section VII concludes the paper. The exoatmospheric missile (M for short) planar dynamics are presented in Fig. 1. The longitudinal thrust generates an acceleration aM , parallel to the fuselage. When deriving the guidance law, we assume aM is constant. In our simulation study, aM is assumed time-varying because the mass is timevarying in practice. The lateral thrust generates a lateral acceleration an. Thus a torque is exerted to M\u2019s nose such that M is rotated, which means the direction of the longitudinal acceleration is also changed. The torque can be expressed as anmla/IM , where m, la, IM denote M\u2019s mass, lever arm and rotational inertia, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001266_1350650117743684-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001266_1350650117743684-Figure1-1.png",
        "caption": "Figure 1. The configuration of an aerostatic spindle.",
        "texts": [
            " To improve the efficiency of CFD modeling, a parametric CFD model is developed to calculate the performance of air bearing, by which the performance of air bearing can be analyzed parametrically. Furthermore, as an application case of proposed optimization algorithm, the performance of an annular thrust air bearing with pocketed orificetype restrictor is optimized. And optimal geometrical parameters for restrictor design are acquired and further verified by experiments, which proves the reliability of proposed optimization algorithm. Parametric CFD model for air bearing performance analysis Figure 1(a) shows the configuration of an aerostatic spindle employed in an ultraprecision machine tool. The shaft is connected to two thrust plates with bolts. Its aerostatic bearing consists of a radial bearing and two thrust bearings. With the support of two thrust air films and a radial air film, the thrust plates and shaft are separated from shaft sleeve. The orifice of restrictor is the inlet of pressurized air. The operating pressure of air spindle is 0.5MPa. The geometrical parameters of thrust air bearing are illustrated in Figure 1(b). The internal and external diameters of air film are D1 and D3, respectively. There are n (orifice number n\u00bc 6) orifices equally spaced on a circumference with diameter of D2. The geometrical parameters of restrictor are listed as following. The diameter of orifice is d, the air film thickness is h, the diameter of chamber is d1, and the depth of chamber is h1. The values of each geometrical parameter are presented in Table 1. Figure 2(a) illustrates the CFD model of thrust air bearing. As the air film is cyclically symmetric, to save calculation time, a basic sector which represents one-nth of air film is built and periodic boundary condition is adopted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001260_pssa.201700468-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001260_pssa.201700468-Figure2-1.png",
        "caption": "Figure 2. a) Photograph and (b) design of the \u03bcPAD.",
        "texts": [
            " The thickness of the paper determines the height of the microchannel which was 100mm in the present work. The prepared sample was then post baked at the 90 C temperature for 5min. Finally, the photoresist was removed from the part of the sample which was exposed to UV light by developing the pattern in SU 8 developer followed by rinsing in iso-propyl alcohol and the desired design (of SU 8 photoresist) was obtained on the paper. The photograph of the prepared paper based microfluidic device as shown in Figure 2(a). A complete design of the paper based microfluidic biosensor using the threeelectrodes is showninFigure2(b).A threeelectrode systemwas prepared on the detection zone of the patterned paper by stencil printing. Graphite was used as both the working and counter electrodes and silver was used as the reference electrode. Phys. Status Solidi A 2017, 1700468 1700468 ( Graphite ink was prepared in house by dispersing pure graphite powder in ethylene glycol and adding polyvinylacetate as binder and subsequently ultrasonicating for 1 h. Silver paste was used for reference electrode after dilution in ethylene glycol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002547_1.4033101-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002547_1.4033101-Figure5-1.png",
        "caption": "Fig. 5 Outer ring cross section with vectors locating the point of contact between ball and outer raceway in the inertial reference frame",
        "texts": [
            " The position vector of the outer raceway is therefore defined as ri R1 \u00bc ri R \u00fe ri RR1 (5) Hereafter, the determination of the overlap and normal force between the ball and outer raceway is consistent with that described by Saheta [11] utilizing the race and azimuth reference frames with the tangential traction force determined in a contact reference frame. The resulting normal and tangential force vectors are added to the total force acting on the center of mass of the outer ring. The moment acting on the outer ring from the ball contact is calculated as a cross product of the position vector and the total force application. A vector locating the point of contact from the outer ring is determined in the inertial reference frame as shown in Fig. 5 ri RS \u00bc ri RR1 \u00fe ri RB \u00fe ri BS (6) The subscript S refers to the point of contact and subscript B refers to the ball. The vector ri RS is used to calculate the moments about the outer ring center of mass due to the normal and tangential forces at the ball and outer raceway contact Ti R \u00bc ri RS Fi tot (7) Ti R is the resulting torque acting on the outer ring from a ball in the inertial frame and Fi tot is the summation of the normal and tangential force vectors in the inertial frame. Ti R is furthermore transformed to the body fixed frame using the transformation in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure3-1.png",
        "caption": "Fig. 3. Continuous medium using DEM and beam model.",
        "texts": [
            " The rings are the most fragile organs in a bearing since they are subjected to cyclic mechanical stresses leading to fatigue or damage such as cracking, pitting or spalling [11] . Therefore the model must be extended to describe deformable rings. In this section, the DEM approach is extended to the description of a continuous medium. In the case of bearing appli- cation, the mechanical behavior of the rings made of steel has to be reproduced. Rings are deformed under mechanical loadings depending on the mounting. The quality of the initial discrete assembly plays an important role in simulating a 2D continuous, homogeneous and isotropic medium with DEM ( Fig. 3 ). To consider a granular assembly as a continuous medium, the rings are descretized by polydisperse cylindrical particles generated with a radius expansion algorithm initially developed by Lubachevsky [37] . The \u201cmicroscopic\u201d assembly must comply with the following properties: \u2022 Isotropic distribution of contact orientations \u2022 Local homogeneous properties: coordination number and local porosity \u2022 Compacity close to 86-87%, according to RCP assumption These considerations involve working with polydispersity given in Table 1 where the adopted discretization offers a compromise between accuracy and calculation time. This compacity called random close packing (RCP) [38] characterizes a maximum density reached with polydisperse particles. In order to reproduce the elasticity behavior of a continuous medium, elastic connections ensure cohesive links between microscopic particles. A beam model based on Euler\u2013Bernoulli theory [39] is introduced and described in Fig. 3 . An particulate lattice model in which the interaction between two cohesive cylindrical particles is modeled by a beam of length L \u03bc, Young\u2019s modulus E \u03bc, cross-section A \u03bc and quadratic moment I \u03bc ( Fig. 3 ). Therefore, the cohesive contacts are maintained by a vector of three generalized components forces acting within the material. The normal component acts as an attractive force, the tangential component allows to resist to the tangential relative displacement and the moment component counteracts the bending motion [39] . l is the thickness of the granular medium corresponding to the length of a rolling element, in the axial direction. Stresses in rings determined subsequently will be assumed to be constant in this direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000169_ext.12016-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000169_ext.12016-Figure7-1.png",
        "caption": "Figure 7 Experimental setup: CAD model.",
        "texts": [],
        "surrounding_texts": [
            "Because of the rotation of the measuring tool, a multichannel telemetry system (Manner Sensortelemetrie, Eschenwasen 20, 78549 Spaichingen, Germany) executes the wireless power supply of the strain gages and the transfer of the measured strain values to the evaluation unit (multichannel receiver). The measuring signal is transmitted by means of induction. All strain data are interpreted as output voltage in a range of \u00b110 V. An MGCplus is used as data acquisition system. The chart in Fig. 6 illustrates the whole experimental environment schematic. Figures 7 (CAD-model) and 8 (real setup) show the installation at the forming machine. Determination of the current tool position To locate the current positions of the measuring points, a linear encoder (Heidenhain MT2571, Dr.-Johannes-Heidenhain-Stra\u00dfe 5, 83301 Traunreut, Germany) and an angle encoder (Heidenhain ROD426, Dr.-Johannes-Heidenhain-Stra\u00dfe 5, 83301 Traunreut, Germany) are installed. The linear encoder provides information about the penetration, whereas the angle encoder captures the tool rotation. A second linear encoder is used to measure the axial shift of the component. Tool and component Experimental Techniques 39 (2015) 28\u201336 \u00a9 2013, Society for Experimental Mechanics 31 kinematics are synchronously processed and recorded with the strain measurements."
        ]
    },
    {
        "image_filename": "designv11_13_0001009_iet-epa.2016.0680-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001009_iet-epa.2016.0680-Figure5-1.png",
        "caption": "Fig. 5 Characteristic map of orientation deviation u with respect to c\u2217 r and T\u2217 e",
        "texts": [
            " 0 always hold for a real IM, the following inequality is always true \u2202 tan u \u2202d . 0 (36) As seen from (27), if d = 0, then u = 0. Therefore, the following law can be obtained sign(d) = sign(u) (37) where sign() is sign function of variable. Its definition is sign(x) = 1 x . 0 0 x = 0 \u22121 x , 0 \u23a7\u23a8 \u23a9 (38) IET Electr. Power Appl., 2017, Vol. 11, Iss. 6, pp. 1105\u20131113 & The Institution of Engineering and Technology 2017 Lines marked with circles and stars indicate experimental results The characteristic map of u with respect to T\u2217 e and c\u2217 r is given in Fig. 5. The scope of c\u2217 r is 0.5\u20131.2 p.u., only including monotonically decreasing part in Fig. 1. L\u2217m is rated value (dashed line in Fig. 1) in standard IFOC controller. In Fig. 5, the red-dashed plane is u = 0 at c\u2217 r = 1 p.u. The plane divides the characteristic map into two parts. In the left part, c\u2217 r is smaller than 1 p.u., Fig. 1 shows that the L\u2217m is smaller than Lm, so d , 0. Then, according to (37), u is smaller than 0, which is consistent with the characteristic map. In the right part, the inequalities c\u2217 r . 1 p.u., L\u2217m . Lm, d . 0 and u . 0 hold. Reference torque also influence the value of u. When T\u2217 e = 0, u = 0 always hold, no matter how much is the c\u2217 r ",
            " 0 (47) a Mutual inductance b Orientation deviation angle c Rotor flux d Torque Solid lines: parameters of motor; dotted lines: parameters of IFOC controller; dashed lines: parameters calculated from (27), (39), (40) 1110 Since DTe and Dcr are zero at d = 0, the following equation is hold sign(d) = sign(Dcr ) = sign(DTe ) (48) The above equation indicates that if mutual inductance used in IFOC controller is larger than actual value of motor, the output torque and flux will be smaller than reference value and vice versa. Combining the magnetic saturation model, it can be deduced that if flux level is higher than rated value, the output torque and flux will be smaller than required value and vice versa. The characteristic maps of cr, Dcr , Te, DTe with respect to c\u2217 r and T\u2217 e are shown in Fig. 6. The dashed planes, marked in Figs. 6a, b and d can be treated as characteristic maps of ideal motor. The planes intersect with maps in the lines c\u2217 r = 1 p.u. Similar with Fig. 5, the planes divide the Dcr and DTe map into two parts. The left part of Dcr and DTe maps are under zero, while the right part of the two are above zero, coinciding with (48). The accuracy of derived equations was investigated by means of simulation and experiment on a 1.1 kw IM with the rated parameters given in Table 1, and magnetic saturation curve shown in Fig. 1. The calculation period of IFOC algorithm is 0.1 ms, so is the pulse width modulation (PWM) wave. The accuracy of equations is demonstrated at different torque and flux-level conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000333_19443994.2014.950986-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000333_19443994.2014.950986-Figure1-1.png",
        "caption": "Fig. 1. Laboratory-scale MBSC-MFC unit.",
        "texts": [
            " Therefore, in this study, a multi-baffled single chamber MFC (MBSC-MFC) using four different wastewaters (cafeteria wastewater [CW], domestic wastewater (influent [IP] and effluent [EP] of primary settling tank in domestic WWTP), and milk processing wastewater [MW]) under continuous flow mode was investigated. In addition, the effect of the OLR and substrate types on power production was analyzed. The MBSC-MFC unit (L \u00d7H \u00d7D; 28 \u00d7 23 \u00d7 2.5 cm) consisted of two separator electrode assemblies (SEAs: SEA1 and SEA2), which shared one anodic compartment (working volume: 650mL). Eight rectangular-type baffles were installed in the anode chamber to enhance the internal fluid flow (Fig. 1). The SEA consisted of an anode, a air-cathode, and a separator. Graphite felt (GF-20-5F, Nippon Carbon, Japan) with a thickness of 5 mm was used as the anode (20 cm [L] \u00d7 15 cm [H]). The air-cathodes were constructed of a stainless steel mesh (SUS 316) with more than 90% purity multi-wall carbon nanotubes (Carbon Nano-material Technology, Korea) as the catalyst and 10 wt.% polytetrafluoroethylene as the diffusion layer and treated based on the study reported by Cheng et al. [10]. Cation exchange membranes (CMI-7000, Membrane International Inc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001942_c5an01200g-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001942_c5an01200g-Figure4-1.png",
        "caption": "Fig. 4 Schematic diagram of individual layers of the device. The device consists of five functional layers: anode/cathode layer, anodic/cathodic paper reservoir, and paper proton exchange membrane. Anolyte and catholyte were trapped in the hydrophilic regions on paper. Several anodic paper reservoir layers were stacked to control the proton/electron travel distance in the MFC system.",
        "texts": [],
        "surrounding_texts": [
            "Fig. 3 describes our assembled 3-D paper based microbial sensing platform. The sensor included four spatially distinct wells of the sensor array. We leveraged the techniques recently demonstrated for paper-based MFC devices,21,24\u201326 and paperbased 3-D culture systems.27\u201329 Each MFC stack contained five functional layers: anode/cathode layers (Au/Cr on 1.6 mmthick polymethyl methacrylate (PMMA)), anodic/cathodic paper reservoir layers (Whatman #1, \u223c180 \u03bcm thick), and a paper-based proton exchange membrane (PEM) (Reynolds Parchment Paper, \u223c50 \u03bcm thick) (Fig. 3 and 4). Each layer was micro-patterned using laser micromachining (Universal Laser System VLS 3.5). The anode and cathode layers were prepared by depositing 200 nm gold on PMMA substrates with 20 nm chrome as the adhesion layer. The layers had metal pads (10 mm diameter) with through-holes in the center for use when we needed to directly introduce anolyte/catholyte into the anodic/cathodic paper reservoirs. Paper reservoirs featuring a hydrophilic chamber with hydrophobic wax boundaries were microfabricated by heat pressing commercially available wax paper onto a filter paper.24 This paper-based platform can be easily directed toward the development of a much higher throughput array."
        ]
    },
    {
        "image_filename": "designv11_13_0001578_s10846-018-0922-5-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001578_s10846-018-0922-5-Figure1-1.png",
        "caption": "Fig. 1 The WMR and its coordinates in two different representations, traversing a predefined geometric path parameterized by \u03c4",
        "texts": [
            " Proof By the Lemma 1, the functions x(t) and y(t) also define the heading angle \u03b8(t) in a C2 class curve; since the curve can be represented by both distance and heading angle function (Lemma 2) The robot input voltage signals u \u2208 R 2, u = [ ur ul ]T impose torque to the wheels. Given the mass m and the moment of inertia J of the robot structure, consider the simplified robot dynamics written as: with matrices defined by: R = \u23a1 \u23a3 Km r Km r Kml 2r \u2212Kml 2r \u23a4 \u23a6 , (6) M = [ m 0 0 J ] , (7) where Km is the motor torque constant, and q\u0308 \u2208 R 2, q\u0308 = [ \u03b3\u0308 \u03b8\u0308 ]T is the acceleration vector. Consider a given path defined mathematically as a function s : [0, 1] \u2192 R 2 such that, q(t) = s(\u03c4 (t)), t \u2208 [0, Tf ], (8) where the robot must traverse a geometric path as illustrated in Fig. 1. The monotonically increasing function \u03c4 : [0, Tf ] \u2192 [0, 1] maps the robot motion time in a normalized interval where \u03c4(0) = 0 e \u03c4(Tf ) = 1. The derivatives of Eq. 8 are the robot velocity and acceleration in the configuration space [7]: q\u0307(t) = s\u2032(\u03c4 )\u03c4\u0307 (t), (9) q\u0308(t) = s\u2032(\u03c4 )\u03c4\u0308 (t) + s\u2032\u2032(\u03c4 )\u03c4\u0307 2(t), (10) where (.)\u2032 refers to the derivatives with respect to \u03c4 . We want to find a velocity profile q\u0307(t), t \u2208 [0, Tf ], such that the robot is driven through a geometric path s in minimum traversal time Tf and energy consumption Et "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001310_etep.2507-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001310_etep.2507-Figure1-1.png",
        "caption": "FIGURE 1 Structure of submersible motor",
        "texts": [
            " According to the first law of thermodynamic, the average temperature of motor is obtained by Gauss\u2010Seidel method, which can realize temperature field and fluid field coupling. Combined with the second law of thermodynamics, it is to improve motor performance by analyzing the entropy generation and the exergy destruction rate of motor and propose some methods to improve the temperature rise of motor. Finally, compared with the prototype temperature test, it shows that the calculated results are close to the test data. This method can provide a certain basis for the motor temperature prediction. The structure of submersible motor is shown as Figure 1. The stator and the rotor are segmented structure, and the segment length is about 0.03 to 0.05 m. The stator is connected by magnetic isolated segment, and the rotor is connected by bearing. The thrust bearing at the top of the motor is mainly used for supporting the rotor, and the motor is filled with the mineral oil, which promotes the fluid flow and enhances cooling. The temperature calculation of the motor involves many disciplines, such as thermodynamics, fluid mechanics, and electromagnetism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001432_s00170-018-1990-1-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001432_s00170-018-1990-1-Figure8-1.png",
        "caption": "Fig. 8 3D FE models with mesh and boundary conditions. a Full-scale model. b Symmetry model. c Quarter model. d Slice model",
        "texts": [
            " The structure deformation, residual stress distribution, and load transfer mechanism may be axisymmetric in some cases, which will be easier to understand with a 3D FE model. It is necessary to build 3D FE models for accurate results. This section discusses the performance of 3D FE models based on force-controlled approach. The E7000 automatic drill- rivet machine (developed by EI company) is operated with force-controlled rivet installation method (https://www. electroimpact.com/). The three-dimensional finite element model contains fullscale model, symmetry model, quarter model, and slice model. Figure 8a shows the conditions of explicit full-scale model. The dimension of the sheet is 50 \u00d7 50 \u00d7 3 mm. Tables 1 and 2 respectively list the material parameters for slug rivet and sheet. The sheet surfaces on the far-end are constrained in radial direction. And, the top and bottom lines of the sheets are constrained at the axial direction. All freedom degrees of the riveting dies and pressure feet are constrained, except the Y-direction. The riveting force is 40,000 N. The clamping force is 500 N. The FE model is generated using C3R8D reduced integration 8-node solid continuum elements",
            " Three deformable bodies, two sheets and a rivet, are defined in the mode. The riveting dies and the pressure feet are defined as rigid bodies. Mesh size of rivet is 0.15 mm. Mesh size of sheet is 0.2 mm for the area in the vicinity of the rivet and 1.2 mm for the region far from the hole. Surface interaction is defined as contact pairs. A friction coefficient of 0.2 is specified for all interactions. The riveting process is simulated in two steps. Step 1 is a loading step. Step 2 is an unloading step. The period of each step is 1 ms. Figure 8b illustrates the symmetry model with its mesh and boundary conditions. The setting of symmetry model is in accord with full-scale model. A X-symmetry plane is added due to symmetric structure. The squeezing force and clamping force are 20,000 and 250 N, respectively, which are half of the forces used in full-scale model. Figure 8c presents the conditions of quarter model. The setting of quarter model is consistent with full-scale model. Two symmetry planes are added. The squeezing force and clamping force are 10,000 and 125 N, respectively, which are a quarter of the forces used in full-scale model. Figure 8d shows the slice model. The sheet is a 15\u00b0 sector portion. The value of this angle is evaluated to avoid radial distortion and stretch ratios of the model [26]. The radius of the sector is 25 mm. The setting of slice model keeps the same with full-scale model. But, in order to prevent the structural deformation in circumferential direction, a new cylindrical coordinate is built to replace the default rectangular coordinate. In Fig. 8d, R is the radial direction, T is the circumferential direction and Z is the height direction. The squeezing force and clamping force are 1666.67 and 20.83 N, respectively, which are one over 24 of the forces employed in full-scale model. If the forces applied in symmetry model, quarter model, and slice model stay the same with the forces used in full-scale model, the deformations of these models would be out of range. The simulation results will be undesired, since the size of these three models is different from the full-scale model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000937_physrevfluids.2.014101-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000937_physrevfluids.2.014101-Figure1-1.png",
        "caption": "FIG. 1. Sketch of the problem.",
        "texts": [
            " When a particle is placed in an electric field, free charges accumulate at its interface due to the difference of electrical conductivity \u03c3 and permittivity \u03b5 between the particle p and suspending s medium [26], R = \u03c3p \u03c3s , S = \u03b5p \u03b5s . (1) A polarization corresponding to a dipole antiparallel to the applied electric field is unfavorable and can lead to rotation. The conditions for continuous rotation are determined from the analysis of the equations for free-charge polarization relaxation and conservation of angular momentum. The evolution of the ith dipole component in a coordinate system centered at the particle [and in the case of an ellipsoid aligned with its axes (see Fig. 1)] is described by [27] dPi dt = \u2212\u03c4\u22121 i [ Pi \u2212 ( \u03c70 i \u2212 \u03c7\u221e i ) Ei ] , (2) where \u03c70 i and \u03c7\u221e i are the low- and high-frequency susceptibilities, respectively, and \u03c4i is the Maxwell-Wagner polarization time along the ith axis of the ellipsoid. The particle rotation with rate is described by Ii d i dt = (P\u0304 \u00d7 E)i \u2212 \u03b1i i, (3) where Ii are the moments of inertia and \u03b1i are the rotational friction coefficients around the ith axis. The electric torque is determined by the total polarization P\u0304i = Pi + \u03c7\u221e i Ei "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001432_s00170-018-1990-1-Figure22-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001432_s00170-018-1990-1-Figure22-1.png",
        "caption": "Fig. 22 The deformation analysis of the automatic drill-rivet machine",
        "texts": [
            " The comparisons of the interference level as well as the residual stress state are somewhat worse than the conditions obtained from force-controlled models. Therefore, the displacement-controlled FE models can only provide a relatively desired simulation results compared with forcecontrolled FE models. The purpose of this paper is to develop some appropriate models for the special applications with a compromise between model size, calculation time, and simulation precision. Some applications of these different FE models are listed as below. Figure 22 illustrates the deformation analysis of automatic drill-rivet machine. Deformation analysis is beneficial to improve the structure design and optimization. To reduce the impact of structural deformation on riveting quality, the (a) Top surface of the sheets. (b) Faying surface of the sheets. (c) Bottom surface of the sheets. Fig. 20 Residual radial stress. a Top surface of the sheets. b Faying surface of the sheets. c Bottom surface of the sheets mapping relationship between slug rivet, aircraft panels, and automatic drill-rivet machine is built by numerical method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000757_j.cirpj.2015.08.005-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000757_j.cirpj.2015.08.005-Figure3-1.png",
        "caption": "Fig. 3. Master on the top and slave on bottom, both of them consist of three magnet groups. Ball transfer unit is placed at the centroid of the equilateral triangle with magnets at its vertices.",
        "texts": [
            " Each group has three magnets at the vertices of an equilateral triangle. These three groups of permanent magnets are located on a circumference 120 degrees apart completing the array (Fig. 2). Similar arrangements are adopted both in the master and slave array. A magnet in the master module array has an opposite magnet in the slave module array, forming a pair. The magnets of each pair are nominally coaxial. The clamping contact point is at the centroid of each group where a ball transfer unit is attached (Fig. 3). A magnet attracts the opposite pole and repels the similar pole. Since these permanent magnets are axially magnetized the N-S-N-S orientation is used to get the highest possible attraction force as shown in Fig. 4. Coaxial magnets located in the master and slave array have their own magnetic attraction forces between them (Fig. 4). However, the other two magnets of the same group have an influence on that force as they are closely located. This was accounted for with the introduction of a force ratio representing the array contact force enhancement by Vokoun and Beleggia [8]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001923_iecon.2014.7048497-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001923_iecon.2014.7048497-Figure1-1.png",
        "caption": "Fig. 1. Active magnetic bearing experiment.",
        "texts": [
            " The designed controllers are implemented in real-time by using a Digital Signal Processor (DSP). Finally, the experimental results are obtained and compared with the conventional lead-lag compensator and the on-board analogue controller while the rotor is stationary as well as while the rotor is in rotation. II. AMB SYSTEM IDENTIFICATION The system under study is an MBC 500 magnetic bearing system. It consists of two active radial magnetic bearings which support both ends of a rotor and levitate the rotor by using electromagnetic force. Fig. 1 illustrates the top and front views of the experimental setup. The rotor shaft has four degrees-of-freedom (4DOF), being translational, namely x1 and x2 in the horizontal direction and y1 and y2 in the vertical direction. The system employs four linear current amplifiers and four decentralised analogue lead compensators to actively control the bearing axes. As AMBs are inherently open-loop unstable, closed-loop system identification must be performed [10], [11]. Various chirp signals are employed as inputs to the system and the outputs from the system are collected while the system is stabilized under on-board analogue controllers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000088_tmag.2013.2291938-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000088_tmag.2013.2291938-Figure2-1.png",
        "caption": "Fig. 2. Rotor, stator, and winding scheme of the SRM 10/8 motor evaluated in this paper.",
        "texts": [
            " Phase A is comprised of four torque component windings (a1\u2013a4). The radial y-component winding is a series of two windings. An advantage of this configuration is that the radial position can be controlled by only two windings. In terms of fail-safe operation, this is a disadvantage, because a single fault in one of the two position control components leads to an overall system failure. This could be prevented by separating them. Our SRM motor topology used in this paper is a SRM 10/8 machine with a pole pair number of \u03bc = 2. As shown in Fig. 2, the SRM 10/8 motor needs two active phases (a and b) for each step to produce torque, whereas the SRM 12/8 motor has just one active phase per step. The pole pair number of \u03bc = 2 and the use of two active phases per step enables shorter flux paths. In contrast to the SRM 12/8 motor in Fig. 1, the five torque windings a\u2013e are not comprised by windings in series. Therefore, the phases a1 and a2 are controlled separately, which means they are in parallel. Another advantage of the SRM 10/8 configuration is that the 0018-9464 \u00a9 2014 IEEE",
            " The radial force components are produced by controlling the 10 phases (a1, b1, c1, d1, e1, a2, b2, c2, d2, and e2) differently. Therefore, the radial force components are produced by all phases and that means a phase fault just decreases the maximum amount of possible radial force, which is not an overall system failure. This topology enables us to use the whole winding space-not just a fraction-to produce radial force. Therefore, much higher radial forces can be produced to counteract gyroscopic torque and linear acceleration in mobile flywheel operation. As shown in Fig. 2, the flux path is short and just a small region of the rotor and stator is magnetized. However, the SRM 12/8 motor in Fig. 1 and all other SRM motors with a pole pair number of \u03bc = 1 have to magnetize the whole stator for each step and the magnetic flux path is going straight through the center of the rotor. Therefore, the SRM 10/8 topology promises less hysteresis losses than conventional SRMs. The SRM 10/8 configuration also makes it possible to have a relatively big rotor compared with the stator, which is good for storing rotational energy, and it increases the overall energy density Nsteps = Nrp \u00b7 Nsp 2 \u00b7 \u03bc "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002036_msf.808.89-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002036_msf.808.89-Figure3-1.png",
        "caption": "Fig. 3 Three dimensional CAD Models of the Total Hip Replacement implants",
        "texts": [
            " In this case study FDM technique has been used to make patterns for investment casting to make a medical implant known as Total Hip Replacement. Two different types of Total Hip Replacement implants have been manufactured from stainless steel 316L. Data regarding their properties has been reported. The different steps of the process are given below: The first step in HIC is preparation of drawing. For this purpose, two different types of drawings of Total Hip Replacement implant were prepared by using Solid works design software. The drawings of the parts are shown in Fig. 3(a) and 3(b). The drawings can be customized according to the size/dimensions required for a specific patient. For some specific medical applications and to make the customized implants, it is also possible to get the drawings from CT scan data of the patient. In order to make the Quick-Cast build style the models were made hollow from inside. The 3D CAD models are saved in .STL format and loaded in the machine software. The machine used is Stratasys uPrintSE as shown in Fig. 2. The suitable orientation of the work piece is fixed and various machine parameters are adjusted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002662_1.4033693-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002662_1.4033693-Figure1-1.png",
        "caption": "Fig. 1 Two kinds of GPEs: (a) external GPE consisting of a sun gear, planet gear, and a carrier and (b) internal GPE consisting of a ring gear, planet gear, and a carrier",
        "texts": [
            " In this study, the principle of equilibrium of torques will be applied to show how the efficiency of an epicyclic GP is related to the efficiency of a conventional GP and its kinematic inversions and to identify the efficiency characteristics of 1 and 2DOF GPEs. General Pair Entity (GPE). The GPE is defined as an entity consisting of a pair of meshing gears and a carrier. A gear or a carrier can be shared by different entities. The EGTs are composed of two kinds of GPEs: the external GPE (planet-sun GP) and the internal GPE (planet-ring GP). A GPE has three basic members that can input or output power, including the carrier as in Fig. 1. A GPE has 2DOFs, i.e., two constraints must be given, otherwise the relationships among the three members are independent. For example, if two of the three members are given as inputs, then the other can be determined as the output. It is also possible to have one input and two outputs, and knowing any two values makes it possible to calculate the third. Planet Speed Ratio. The term \u201cplanet speed ratio\u201d will be used here to denote the relative speed ratio between any two links of a GPE with respect to the carrier [26]",
            " The planet gear ratio Np;q can also be written for any GP p and q as Np;q \u00bc 6 Zp Zq (2) where Zp and Zq denote the number of teeth on gears p and q, respectively, with the plus sign corresponding to the internal GP and minus sign to the external GP, respectively. Speed Ratio. The term \u201cspeed ratio\u201d is used to denote the speed ratio between two links of the GPE with respect to the third. Let the symbol Rv w;u denote the speed ratio between links u and w with respect to link v. Then, the speed ratio Rv w;u is Rv w;u \u00bc xw xv xu xv \u00bc Np;w Np;v Np;u Np;v (3) Potential Power Efficiency and Power Loss. For the GPE shown in Fig. 1, let Tp, Tq, and Tc denote external torques acting on links p, q (sun gear or ring gear), and c, respectively. The relation between Tp, Tq, and Tc is Tp \u00fe Tq \u00fe Tc \u00bc 0 (4) Under constant-speed operation and by taking friction power losses L in consideration, the power balance of a GPE is Tpxp \u00fe Tqxq \u00fe Tcxc \u00fe L \u00bc 0 (5) By multiplying Eq. (4) by xj, where j is any link in the GPE, and subtracting the product from Eq. (5), we get Tp\u00f0xp xj\u00de \u00fe Tq\u00f0xq xj\u00de \u00fe Tc\u00f0xc xj\u00de \u00fe L \u00bc 0 (6) Equations (5) and (6) lead to the important observation that the power loss of any GPE is frame-independent and that the torques can be calculated in any moving reference frame",
            " (24) yields gq p c\u00f0 \u00de \u00bc Np;q gGP Np;q 1 (25) Applying the six power flow arrangements for any 1DOF GPE, six efficiency relations can be derived (Table 1). Pennestr \u0131 et al. [11] embodied a more extended table of mechanical efficiencies of 1DOF gear train kinematic inversions. Indeed, due to the absence of a specific definition of the gear ratio R (Nj;i), the extended table reported by Pennestr \u0131 et al. [11] contains a redundancy in the efficiency formulas of 1DOF GPEs. A more restricted definition for the gear ratio is used in this paper and is called the planet gear ratio, Np;q \u00bc 6Zp=Zq. For the external GPE shown in Fig. 1(a), Np;q\u00f0\u00bc Zp=Zsun gear\u00de is always smaller than zero (Np;s < 0), while for the internal GPE shown in Fig. 1(b), Np;q\u00f0\u00bc Zp=Zring gear\u00de has always positive value but since the ring gear is always larger than the planet gear, it is always smaller than 1 (0 < Np;r < 1). By virtue of the concept of the planet gear ratio, the number of gear ratio ranges for any GPE is reduced to only two ranges. Although the current efficiency formulas are less than what is known in previous literature, they cover all the possible cases. Preliminary Discussion. Apparently, beginning with the paper by Radzimovsky [6], formulas susceptible to wrong application were used in the literature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003412_15397734.2018.1457446-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003412_15397734.2018.1457446-Figure4-1.png",
        "caption": "Figure 4. The dynamic model of the second model with springs and dampers.",
        "texts": [
            " The gravitational acceleration is taken 10 m/s2.An initiallinear velocity of 60 km/h and the corresponding initial rotational velocity is imposed onthe wheel to ensure its rollingmotion on the rail at the start of the analysis. This velocity of 60 km/h is maintained and applied to the central axis of the wheel as a boundary condition to avoid the reduction of the train speed.A frictional force is generated between the wheel and the rail due to this boundary conditionwhich assuresthe rolling motion of the wheel. Figure 4 illustrates the dynamic model of the second model and Table 2 presents the input parameters.According to common practices, wheel flats should be restricted to a length of 60 mm and a depth of 0.9-1.4 mm (Vyas and Gupta 2006). We have chosen a flat of 0.7mm depth and 51.3mm long. Table 3 lists the dynamic properties of the rail foundation for seven different cases. Generally, the results of the second model arebased on the properties of the second case in Table 3. In the case other properties are used, the case number will be mentioned"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001022_tia.2017.2683439-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001022_tia.2017.2683439-Figure8-1.png",
        "caption": "Fig. 8. The analyzed geometries of the stator laminations. (a) the geometry of produced motor prototypes, (b) & (c) alternative geometries.",
        "texts": [
            " In (8), P is iron loss in stator core, Bi is module of the flux density vector existing at i-th mesh element, pi(Bi) is specific iron loss at Bi flux density (determined with the SST system for axial magnetization), mi is mass of i-th mesh element.    n i iii mBpP 1 )( (8) Since the teeth iron losses represent the main components of the total core losses, the omission of the rotational magnetization (occurring in a small part of the stator yoke only) results in a negligible error. To investigate the influence of the stator lamination geometry on the BF factor, the authors executed a simulation of three lamination geometries shown in Fig.8. The (a) and (b) cases concern a situation where the authors left the same outer and inner dimensions of the stator lamination (outer diameter 100 mm, inner diameter 56 mm), and the same area of the stator slots which respect to produced motor. The third case (c) concerns a situation where the outer diameter of the stator and rotor was increased of 25 mm, and the width of the stator teeth was such that undamaged areas of the material could appear. This outer diameter corresponds to the motor with a nominal power of 1",
            " In the computation of the iron losses in the damaged zones, As discussed in [11], the losses inside the damaged area change if the magnetic flux flows parallel to the \u2018green\u2019 material with respect to the situation in which the magnetic flux flows perpendicularly from the damaged area to the undamaged one. The building factor BFg of the stator core, resulting from the stator core geometry, was computed in accordance to (9): sg spd g P P BF  (9) where Psg are the iron losses in the \u201cgreen\u201d stator core, Pspd are the iron losses in the \u2018partially damaged\u2019 stator core. At rated voltage and rated frequency (50 Hz), for the motor having geometry presented in Fig.8.a, the calculated iron losses in the stator core were 8.5 W and 11.8 W for \u2018green\u2019 material and for \u2018partially damaged\u2019 material, respectively. The difference between these two values agrees with the difference measured in the same condition during the tests (referring to the 550 W motors, Fig.12). On the basis of the FEM results, the calculated BFg building factor of the nonannealed stator core is equal to 1.4. A comparison of the BF factor designated for other motor geometry lamination is included in Table II"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000267_j.measurement.2013.12.012-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000267_j.measurement.2013.12.012-Figure6-1.png",
        "caption": "Fig. 6. Experimental system.",
        "texts": [
            " Since that the bicoherence is a symmetrical matrix to the main diagonal, denoted as \u00bdM \u00bc b0W ;T\u00f0f1; f2\u00de, it can be decomposed as \u00bdM \u00bdVk \u00bc kk\u00bdVk \u00f015\u00de where kk is the eigenvalue, [Vk] is the eigenvetor corresponding to kk. The value of eigenvalue is proportional to the amount of correlation in the direction of their associated eigenvectors, and the maximal eigenvalue linearly increases with the increase of the correlation strength in a cluster. Therefore, the maximal eigenvalue may be defined as a global phase coupling index for bicoherence. Experiments are conducted on the machinery fault simulator (MFS) from SpectraQuest, which is illustrated in Fig. 6. In the fault diagnosis experiment of roller element bearing, the MFS is composed of an AC motor, a coupling, two identical support bearings, a bearing load and a shaft. In this experiment, three bearing health conditions are considered, which includes inner race defect, ball defect and normal bearing. Since vibration-based analysis is one of the principal tools for mechanical fault diagnosis [23\u2013 25], the vibration is measured by the accelerometers mounted on the bearing house at the motor speed 3000 r/min"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000943_iet-cta.2011.0611-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000943_iet-cta.2011.0611-Figure1-1.png",
        "caption": "Fig. 1 Flexible joint electric drive system",
        "texts": [
            " Then, the dynamics of ea(t) is governed by e\u0307a = Naea + TaE\u0304a(d\u0304 + \u03041(x + G1\u03c6)) \u2212 TaF2a\u03d22(Haxa, Haea) = Naea + TaE\u0304a(d\u0304 + \u03041(x + G1\u03c6)) \u2212 F2a\u03d22(Haxa, Haea) This error dynamical system has the same structure as (6). So, according to Theorem 1, if there exist matrices PT a = Pa > 0, Ma and Ka satisfying \u23a7\u23a8 \u23a9 TaE\u0304a = (In+\u03c5 \u2212 MaCa)E\u0304a = 0 PaF2a = H T a N T a Pa + PaNa \u2212 \u03bcH T a Ha < 0 (20) then the estimation error ea(t) converges globally to 0. Furthermore, since the condition (20) has the same form as the one in (9), the parameters Ma and Ka can also be designed according to Theorem 2 and Algorithm 1. Example 1: The non-linear UIO is designed for a flexible joint electric drive system depicted in Fig. 1. The load and the DC motor are connected by a torsional spring. The dynamic equations of the system are given by \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u03b8\u0307l = \u03c9l \u03c9\u0307l = \u03ba Jl (\u03b8m \u2212 \u03b8l) \u2212 Kv Jl \u03c82(\u03c9l) \u03b8\u0307m = \u03c9m \u03c9\u0307m = \u03ba Jm (\u03b8l \u2212 \u03b8m) \u2212 Kv Jm (\u03c9m + \u03c81(\u03c9m)) + K\u03c4 Jm (u + d) (21) where \u03b8l(t) and \u03c9l(t) are the load position and velocity, respectively. \u03b8m(t) and \u03c9m(t) are the motor position and velocity, respectively. The quantities \u03b8m(t), \u03b8l(t) and (\u03c9m(t) \u2212 \u03c9l(t)) are measured. The non-linearities in the system have the form \u03c81(\u03c9m) = { 1 \u2212 e\u2212\u03c9m , \u03c9m 0 \u22121 + e\u03c9m , \u03c9m < 0 \u03c82(\u03c9l) = \u03c9l + \u03c93 l The constants in (21) are given as follows"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000397_j.bios.2013.11.045-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000397_j.bios.2013.11.045-Figure1-1.png",
        "caption": "Fig. 1. Section of the large-volume wall-jet flow cell used. It comprises a copper wire (a), a technical PVC cylinder (b), a \u201csubstrate\u201d electrode (c) an electrically conducting composite), and a gelatin B coating layer (d). The outlet of the FIA system (e) polyether ether ketone HPLC tubing) is directed perpendicularly to the sensor surface, at a 100 \u03bcmdistance. \u201cmV\u201d denotes a high impedance voltmeter, \u201cf\u201d is a reference electrode and \u201cg\u201d is a data station. The figure is not drawn to scale. See experimental for details.",
        "texts": [
            " To avoid such effects at the end of the square concentration pulse of 80 s duration, the injector was switched from inject to load after 80 s (well before the sample loop volume was totally emptied). This results in a sharp pulse with negligible broadening as well at the start as at the end of the pulse. The eluent was 10.0 mM 2-(N-morpholino)ethanesulfonic acid (MES), pH 7.0. The column outlet was directed perpendicularly towards the sensitive membrane of the coated wire electrode in a \u201cwall-jet\u201d flow cell (see, Fig. 1). The distance from the LC tubing outlet to the electrode was 100 mm. This home-made flow-cell is fully described in an earlier publication form the group (Daems et al., 2013). The membrane potential was measured against an Orion 800500Rosss reference electrode (Ag/AgCl) using a high impedance (1013 \u03a9) homemade amplifier. The detection signals were recorded on a data station composed of a PC equipped with a 6013 NI DA converter and LabVIEW 7 (National Instruments, US) based software. The overall RC time constant of the high impedance amplifier plus data station was set to 0",
            " Basic drugs promazine, lidocaine and ritodrine were obtained from Sigma-Aldrich (Bornem, Belgium), dopamine (DA) was available from Fluka. To dissolve the drugs, small amounts of ethanol (Fluka, Analytical grade) were used. The coupling agents 1-ethyl-3(3-dimethylaminopropyl)carbodiimide (EDC) and N-hydroxysuccinimide (NHS) were obtained from SigmaAldrich. The oligonucleotides (aptamers and DNA) were synthesized by Integrated DNA Technologies (IDT, Leuven, Belgium). An important component of the potentiometric sensor is the ionically conductive coating layer, labeled \u201cd\u201d in Fig. 1, which is doped with the bait biomolecule. In the present study, gelatin B is used as a coating material. It is a proteinaceous collagen derived material, well-known for its robust film-forming properties (Garies and Schreiber (2007)). In potentiometric sensors, this film can be used as a thick layer (mm size range), a thin layer (nm size range), or it can even be omitted (gold surface). SPR does not have this choice: only nm bait layer approaches are possible, and the choice of surface materials is limited (gold being the standard material)",
            " The generated potential is directly related to the Gibbs free energy of interaction of the analyte with the surface (Daems et al., 2013; Nagels, 2004) as predicted by the well-known \u0394G\u00bc nFE equation. Herein, \u0394G is the difference in Gibbs free energy of the analyte between the eluent buffer phase and the sensor coating surface. E is the change in potential which results from this interaction. Gelatin B has been used previously (De Wael et al., 2012b) in amperometric sensor applications. It has a good adhesion to our substrate electrode material (Fig. 1c), which is an electrically conducting carbon composite. The covalent binding of aminated and fluorescently labeled oligonucleotides (see Section 2.1.3) to the Gelatin B (which contains carboxyl groups), was confirmed with confocal microscopy. In further experiments, the fluorescent label was omitted. As gelatin B is a proteinaceous material which contains functional groups such as carboxyl groups, covalent coupling of biomolecules with many different known coupling methods is possible. Therefore, doping of this (or an equivalent) ionically conducting coating material offers a good method to incorporate biomolecules"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000439_s10846-014-0063-4-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000439_s10846-014-0063-4-Figure1-1.png",
        "caption": "Fig. 1 The aircraft are spaced between their neighbors by \u0394C , and the circle has radius RF",
        "texts": [
            " Typical examples are: multiple unmanned aircraft coordination [1, 2], task assignment [3, 4], patrolling or surveillance [5\u20137], trajectories generation [8\u201311], searching [12\u201314], target tracking [15\u201317], navigation planning [18, 19], formation [20\u201322] and source seeking [23]. The rising interest on unmanned aircraft squadrons is related to their important proprieties: a distributed system is usually more robust [24] and cheap [25] than a centralized one and the agents make a better use of sensors [26], since the sensors can be shared by a network. Unmanned aerial vehicles can execute important tasks, such as monitoring. This can be made by the squadrons by assuming a circular formation, as shown in Fig. 1, where \u0394C is the arc between two consecutive aircraft and RF is the circular formation radius. Usually, circular formation is not used during take off and landing, and other formation may be necessary. One attractive formation that may be used for landing and take off is the longitudinal formation showed in Fig. 2, where the aircraft are equally spaced by \u0394L. Therefore, an unmanned aircraft squadron that has to execute monitoring, take off and landing tasks will eventually need to change between longitudinal and circular formations. Thus, this is the problem considered here: to implement an unmanned aircraft squadron behavior, by using algorithms, that makes the squadron leave the circular formation shown in Fig. 1 and assume the longitudinal formation shown in Fig. 2, and vice-versa. Different names are given to similar problems in the literature. Some examples are: formation control [27], circular formation [28] and formation reconfiguration [29]. These problems have been investigated in the literature by considering different agents and different approaches. Some examples of agents are: unmanned aerial vehicles [30\u201334], spacecrafts [35\u2013 37], satellites [38\u201340], mobile robots [41\u201343] or simply agents [27\u201329]",
            " In our previous work [53], a control scheme of autonomous aircraft squadrons is proposed by using 3D maneuvers and the virtual structure approach. However, a test trajectory is statically specified, and the formation reconfiguration problem is not treated. After an exhaustive investigation in the literature, it was not found any reference to the problem considered here. Hence, the main contributions are: 1. A distributed algorithm that makes an unmanned aircraft squadron leave the circular formation shown in Fig. 1 and assume the longitudinal formation shown in Fig. 2 and vice-versa. The theoretical analysis ensures its efficiency and the algorithm correction is proved. 2. A new approach based on segments is used to design the proposed algorithms. The methodology addresses all the important aspects previously described, such as the building of efficient algorithms, the formation reconfiguration capacity and the verification of reconfiguration stability. Besides, it is implemented in a distributed way, thereby contributing to the system robustness"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000053_1464419314566086-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000053_1464419314566086-Figure6-1.png",
        "caption": "Figure 6. Roller raceway elastic contact in the SRB with roller error.",
        "texts": [
            " Substituting equations (25) and (26) into the static equilibrium equations (23) and (24), two nonlinear equations with the two unknowns r and \u2019 can be obtained, which can be solved with the Newton\u2013 Raphson iterative method.34 The roller load value and axis position of the SRB can be calculated when the rollers are in any position by changing the value of . Elastic deformation analysis of roller raceway with roller error In actual SRBs, the rollers generally have some diameter error, which will affect roller load distribution and bearing axis orbit.22 Figure 6 shows No. 1 roller has a negative diameter error. D0 is the nominal diameter, and D1 is the actual diameter of No. 1 roller. The actual elastic contact deformation of No. 1 roller and the raceways can be determined by the difference of 1 1 and the roller diameter error. If 1 1 5 \u00f0D0 D1\u00de, it indicates that the elastic contact deformation of No. 1 roller and the raceways is smaller than the roller diameter error and there is no elastic contact between them (shown in Figure 6(a)). Then the actual elastic deformation \u00f0 01 1 \u00de of No. 1 roller and the raceways is equal to 0, i.e., 01 1 \u00bc 0. If 1 1 4 \u00f0D0 D1\u00de, it indicates that the elastic contact deformation of No. 1 roller and the raceways is at Kungl Tekniska Hogskolan / Royal Institute of Technology on September 10, 2015pik.sagepub.comDownloaded from larger than the roller diameter error, and No. 1 roller and raceways are elastically contacting (shown in Figure 6(a)). Then, the actual elastic deformation of No. 1 roller and the raceways can be expressed as 01 1 \u00bc 1 1 \u00f0D0 D1\u00de \u00f029\u00de Substituting equation (25) into equation (29), and noting that the elastic deformation of roller raceway should be nonnegative, the actual elastic deformation of roller raceway in the first row at different azimuth angle can be calculated by 01 j \u00bc max r cos\u00f0 1j \u00fe \u00de cos Pd 2 \u00f0D0 Dj\u00de , 0 \u00f030\u00de Similarly, the actual elastic deformation of roller raceway in the second row at different azimuth angle can be obtained 02 j \u00bc max r cos\u00f0 2j \u00fe \u00de cos Pd 2 \u00f0D0 Dj\u00de , 0 \u00f031\u00de where, Dj is the actual diameter of the jth roller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000081_12.2084733-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000081_12.2084733-Figure3-1.png",
        "caption": "Figure 3, Corrugated actuator  2,2  BA in (a) initial state and (b) deformed under applied voltage.",
        "texts": [
            " 2VM   (7)   bK rrY n i iii     1 1 )log()log(  (8)   K rr rr rrY rr Yrr n i n i ii iiii iii n i ii ii i n i ii                                         1 1 1 01 1 1 1 0 1 1 )(2 )( )log()log()(    (9)                          n i iiii n i iii n i ii rrYYrrYrrK 1 1 1 22 1 2 1 1 loglog 2 1 (10) With the change in angle subtended by the curved unimorph known (7), a relationship between voltage and length change can be derived for the corrugated actuator. Consider the corrugated actuator with A curved segments and B straight segments, in initial and deformed states shown in Figure 3. The change in length of the actuator (11) can be expressed as the sum of the differences in the projected length of the curved segments a and straight segments b . The subscripts d and i denote the deformed and initial states, respectively. )()( idid bbBaaAL  (11) The change in projected length of the straight segment (12) depends on its half-length l and the subtended angle of the curved segment in its initial and deformed state.                     2 cos 2 cos2  lbb id (12) The change in projected length of the curved segment (13) is equal to the change in its chord length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002660_978-3-319-06590-8_48-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002660_978-3-319-06590-8_48-Figure4-1.png",
        "caption": "Fig. 4 SpectraQuest Bearing Prognostics Simulator. Left general view; right floor plan [7]",
        "texts": [
            " Because the step response is dominated by lower frequency content than the impulse response, neglecting this delay would have the effect of underestimating the spall size, and the error would increase with speed as the delay became a larger proportion of the time to impact. Clearly, however, this potential error source does not account for the positive correlation between spall size estimate and speed exhibited in Fig. 3. The test rig used in this study was a Bearing Prognostics Simulator (BPS) provided by SpectraQuest Inc [7], and is shown in Fig. 4. It can be seen in the floor plan that the shaft weight is taken by two support bearings, creating a cantilever arrangement for the test bearing, which has a floating housing through which a purely radial force can be applied. The vibration was measured directly on the test bearing housing using a Br\u00fcel and Kj\u00e6r 4394 IEPE-type accelerometer, which was stud-mounted horizontally to measure vibration in the direction of the applied load. Testing was also conducted with an accelerometer mounted vertically on top of the housing, but this was found to be far less sensitive to the step response from the entry event"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002889_s40436-016-0158-1-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002889_s40436-016-0158-1-Figure6-1.png",
        "caption": "Fig. 6 The thermal deformation of the bearings",
        "texts": [
            " Front and rear bearings are the parts of high temperature and large deformation in the motorized spindle. The thermal deformation of the ends of the motorized spindle is larger due to the deformation accumulation. The cooling water passage of the motor takes away a portion of the heat, thus the thermal deformation is relatively small. As shown in Fig. 5, the main thermal deformation direction of the stator and the rotor is radial deformation, and the deformation is gradually accumulated from center to edge. The front and rear bearing thermal deformation is shown in Fig. 6. The differences of the front bearings inner ring and the outer ring thermal deformation are around 4 lm and 7 lm, respectively, and the differences of the rear bearings are around 2 lm and 4 lm. The front bearings thermal deformations are larger than the rear bearings both in radial and in axial directions. Commonly used in engineering is the first two order natural frequency of the rotor part. For the static rotor, the axial preload is 60 N, and the simulation result the first two order natural frequency is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001537_0954406218791636-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001537_0954406218791636-Figure1-1.png",
        "caption": "Figure 1. (a) The adjustable bearing in its \u2018default adjustment\u2019 as a plain two-sector cylindrical bearing. Rotor centre appears to coincide with the bearing centre. (b) Effective surfaces of hydrodynamic lubrication of a partial-arc bearing of bore radius Rb.",
        "texts": [
            " The various bearing configurations that the adjustable bearing can obtain have been thoroughly investigated in the past from the pioneers on hydrodynamic lubrication and this paper is not an insight into the hydrodynamic lubrication principles under which the various bearing configurations operate. Adjustable bearing configuration and functionality The principle of operation of the proposed bearing with two adjustable pads is on the relative displacement of the bearing pads with respect to the bearing centre. Bearing centre is defined as the centre of the bearing shell (not the pads). In Figure 1(a), the bearing centre coincides with the centred rotor. Nevertheless, in Figure 1(a), each of the pad centre coincides also with the bearing centre, but as it will be presented in continue, each of the pad centres may not always coincide to the bearing centre according to the adjustment. In Figure 1(a), the bearing pads are fixed in the default position and they form a 2-partial-sector cylindrical bearing. The partial sectors of the inner bearing lubricated surfaces are arcs of the same circle of radius Rb, and each of them extents at an angle of p \u00bc 60 (\u2018p\u2019 stands for \u2018pad\u2019) as in many industrial applications.3,35 The radial clearance of the bearing is defined as cb \u00bc Rb R (\u2018b\u2019 stands for \u2018bore\u2019) where R is the journal\u2019s radius, meaning that the radius of the largest journal that can fit in the bearing is Rb. This notation will be valid in the entire paper. In the default adjustment, the bearing centre coincides to the pads\u2019 centre as in Figure 1(a) the bearing centre. The rotor centre would execute any trajectory inside the clearance circle of the pads as in a conventional bearing. Having defined the \u2018default adjustment\u2019 of the bearing, four different adjustments are described in continue, corresponding to different functions of the proposed bearing. The mechanism for adjustment is not the subject of this paper as it is a mechatronic mechanism under development. The \u2018adjustment mechanism\u2019 shown in figures consists of servomotor(s) and potentially piezoelectric elements that enable the displacements of pads, applying quasi-static force that is mechanically amplified and in vertical direction may exceed the 300 kN in the case of a turbine of medium size (100\u2013 200MW)",
            " To achieve this, a geometric preload is introduced to the pads, with respect to the bearing centre, by displacing each pad in the vertical direction (both approaching the journal) by a displacement c, see Figure 3. In this way, the effective eccentricity of the established fluid films is getting increased and thus, the stability characteristics are improved.1,10 The centres of the pads lay in a common circle of radius c, called \u2018preload circle\u2019. The largest journal that would fit in the bearing has a radius of Rb \u00bc R\u00fe c0b, with c0b \u00bc cb c, where cb is the clearance of the bearing with its pads concentric to the bearing centre (see Figure 1(a)). The radius of the pad measured from its centre is then Rp \u00bc R\u00fe c0b \u00fe c or Rp \u00bc R\u00fe cb. With the pad clearance (called also \u2018machined pad clearance\u2019) to be cp \u00bc Rp R, and the bearing clearance (called also \u2018assembled bearing clearance\u2019) to be in this adjustment c0b \u00bc Rb R, the formula m \u00bc 1 c0b=cb yields the geometric preload (or just \u2018preload\u2019) of the bearing. In the current application, m \u00bc c=cb and a suggested initial preload value to check the bearing performance is m \u00bc 0:5. In general, the higher the preload m is, the more stable the bearing is, as the bearing obtains more sharp lemon shape",
            " To avoid confusion, it should be stated that in a conventional lemon bore bearing, the machined pad radius is defined as Rp \u00bc R\u00fe cp, with cp (machined pad clearance) to be different (higher) than cb (assembled bearing clearance), and not equal as in the current example. In the current example, the bearing profile obtains its lemon configuration due to the displacement of the pads and not to the form of the pad. Under this latest consideration, in the current example, the lemon bore adjustment would require a relatively high assembled bearing clearance cb of the reference bearing (Figure 1(a)) so as the displaced (by c) arc to compose a relatively lemon shape as in Figure 3. In other words, the current possibility of adjustment would rather yield a smaller/higher bearing clearance than a lemon bearing unless the clearance is relatively high, e.g. >2.5% of journal radius. Offset bore (offset halves) adjustment Very similarly to the \u2018lemon-bore adjustment\u2019, the \u2018offset-bore adjustment\u2019 considers again a relative displacement of the two bearing pads, but unlike the lemon bore adjustment, in the offset bore adjustment, the bearing pads will be displaced only in horizontal direction, see Figure 4",
            " The reference rotordynamic evaluation is presented in section Reference rotordynamic evaluation including key parameters of the design and performance for the medium size turbine. The initial rotordynamic evaluations used as reference are presented in this section. The schematic representation of the medium size turbine-generator is given in Figure 6; this is a shaft line consisting of three rotors and four bearings. Note that bearings #3 and #4 have been lifted to align the shaft line per GE\u2019s standards for minimum bending moment in the rigid couplings. All four bearings are plain 2-partial arc bearings (equal to adjustable bearing in the default adjustment, see Figure 1). The turbine-generator shaft line executes a virtual transient run-up from 0 to 3000 r/min in a time of 30 s (in real application this duration may be more than 10min) with a constant rotating acceleration at the rate of 100 r/min/s and then continues in steady-state operation of 3000 r/min rated speed for another 20 s to produce the steady-state response that the system is supposed to produce running on site. The basic dimensions of the rotor-bearing system are described in Table 1 while the bearing dimensions and specific loads are described in Table 2",
            " However, the oil film impedance forces FB X and FB Y (see Figure 7), applied from the oil film to the pedestal and the rotor, are always nonlinear functions of the displacement and velocity of the respective journal and pedestal. In the virtual run-up of the system and the steadystate operation at 3000 r/min, each journal will be performing trajectories inside the bearing clearance as shown in Figure 8. The conventional 2-arc 60 bearing is considered in this calculation (equal to the configuration of Figure 1). In each of the graphs of Figure 8 there are indications for the respective clearance circle, the respective bearing bore centre (showed as \u2018bearing centre\u2019), and the respective casing centre that is the point that the casing stator centreline is supposed to pass through. Note that the initial condition for initiating the run-up from zero speed sets each journal centre in the centre of the respective bearing. Starting from zero speed, each journal follows a very fast downward motion before the speed raises approximately above 100 r/min and the hydrodynamic lubrication principle starts taking place",
            " Therefore, the possibility of adjusting the journal centre according to the achieved clearance distribution is a benefit form the proposed bearing and this was discussed already in the section Eccentric rotor whirling. The operating journal eccentricity affects the bearing performance intensively. It is suggested that bearings should operate in an eccentricity of 45\u201365% with respect to radial clearance to achieve stable operation and keep power losses and temperature rise in reasonable values. The power loss for the conventional and the adjusted (default adjustment, see Figure 1) operation was found in the section Eccentric rotor whirling to be as in Table 5. The bearing power loss PLOSS W\u00f0 \u00de due to frictional forces can be evaluated in a fluid film quite accurately using equation (1). PLOSS \u00bc R T\u00f0 \u00de R \u00f0 \u00de \u00f01\u00de R m\u00f0 \u00de is the journal radius, rad=s\u00f0 \u00de is the journal\u2019s rotating speed, and T N\u00f0 \u00de is the resulting shearing force acting on the lubricated surface of the journal, meaning from the starting angle 1 to the ending angle 2 of the developed fluid film wedge. The cavitation of the lubricant is considered with the Gu\u0308mpel (Half Sommerfeld) boundary condition, meaning that only the positive pressure P x, \u00f0 \u00de is accounted in the evaluation of formula in equation (2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure1-1.png",
        "caption": "Fig. 1. Calculating grid for the test of turbulence models.",
        "texts": [
            " For instance, k\u2013e model assumes, that the turbulent viscosity is associated with the kinetic energy and diffusion of turbulence, while the k\u2013x \u2013 turbulent viscosity is related to the kinetic energy and frequency of turbulence. Few of mentioned turbulence models are in addition to the models k\u2013e and k\u2013x (BSL model combines the advantages of k\u2013e and Wilcox) [5]. In order to test the turbulent models, the numerical model (a cylinder in the channel) was developed. Dimensions of the cylinder: diameter is 0.04 m, length is 0.023 m. Dimensions of the channel: width is 0.2 m, length is 0.023 m, height is 0.5 m. The air flows through the channel, flow velocity is 10 m/s. Fig. 1 shows the calculating grid made by the ANSYS CFX code. Five points on the surface of cylinder are marked in Fig. 2. Points are located clockwise. Test results are presented at Table 1. As can be seen, turbulence modeling results are fairly good and confirm to the experimental data. k\u2013e model was selected for the further calculations, as more suitable for modeling turbulent flow at the small flow velocities. where lt \u2013 the eddy viscosity or turbulent viscosity, which must be modeled using k\u2013e model, Cl is a constant equal to 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure6-1.png",
        "caption": "Fig. 6. Fillets of the rack-cutter: a) boundary fillets; and b) real fillets.",
        "texts": [
            " The equation of the radius of the curvature of the curve \u03be is obtained in an analogous way as the curve \u03b7, with the use of Eq. (6). In this case the first and second derivatives of X\u03be and Y\u03be to \u03c6, after differentiating Eq. (10), are defined by the equations. _X \u03be \u00bc r\u2212rb cos\u03c6; \u20acX \u03be \u00bc rb sin\u03c6; _Y \u03be \u00bc \u2212rb sin\u03c6; \u20acY \u03be \u00bc \u2212rb cos\u03c6; \u00f012\u00de e curvature radius \u03c1\u03be=\u03c1\u03be(\u03c6) of the curve \u03be is obtained by the formula \u03c1\u03be \u00bc r2\u22122r rb cos\u03c6\u00fe r2b 3=2 r2b\u2212r rb cos\u03c6 \u00bc mz 1\u22122 cos\u03b1 cos\u03c6\u00fe cos2\u03b1 3=2 2 cos\u03b1 cos\u03b1\u2212 cos\u03c6\u00f0 \u00de : \u00f013\u00de In order to clarify why the rack-cutter fillet undercuts the gear teeth, in Fig. 6a the two boundary curves \u03b7 and \u03be are drawn simultaneously in the current position, where their common contact point coincides with the starting point A of the line of action AB (the position where \u03c6=\u03b1). In this position the radial line OE (Fig. 6a), representing simultaneously a rectilinear profile of the rack-cutter, crosses the curve \u03be in its inflection point, which, in this case, coincides with point A and simultaneously appears as a contact point of OE with the curve \u03b7. From Fig. 6b it becomes clear that if the rack-cutter fillet is an arc of a small radius \u03c11 there exists no undercutting \u2014 type IIa and type IIb, because in this case the arc AF1 lies on the internal side of curves \u03b7 and \u03be. When the rack-cutter fillet is positioned between both curves \u03b7 and \u03be (the arc AF2 of a radius \u03c12>\u03c11) an undercutting\u2014 type IIa appears, and when the rack-cutter fillet is placed between the curve \u03be and the line OE (the arc AF3 of a radius \u03c13>\u03c12) besides an undercutting \u2014 type IIa, an undercutting \u2014 type IIb is derived",
            " It is immediately seen that by increasing the number of teeth z, the depths h\u03b7 and h\u03be of the respective boundary areas increase as well. The same depths, with one and the same number of teeth, decrease when increasing the pressure angle \u03b1. In the case where the rack-cutter fillet is profiled on an arc from a circle, the following boundary condition is defined: the undercutting \u2014 type II (IIa and IIb) is avoided if the radius of the fillet is smaller or equal to the radius of the curve \u03c1\u03b7,A (Fig. 7) of the boundary fillet \u2014 type IIa (curve \u03b7) in point A. Since in point A of the curve \u03b7 (Fig. 6) the value of the angular parameter is \u03c6=\u03b1, for the radius of the curve \u03c1\u03b7,A in this point according to Eq. (9), it is obtained \u03c1\u03b7;A \u00bc mz sin\u03b1 : \u00f017\u00de Then the boundary condition for a non-undercutting of type II, when the fillet curve of the rack-cutter is an arc \u043ef a circle of a radius \u03c1, is defined by the inequality \u03c1\u2264\u03c1max \u00bc \u03c1\u03b7;A \u00f018\u00de finally written as the following type(19) and is \u00f019\u00de where \u03c1*=\u03c1/m is a coefficient of the radius of the circle, on which the rack-cutter fillet is profiled. The condition (19) shows that the undercutting \u2014 type II depends only on the number of teeth z of the gear and the profile angle \u03b1 of the rack-cutter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000616_aim.2012.6265962-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000616_aim.2012.6265962-Figure3-1.png",
        "caption": "Fig. 3. Gaits including contacting-foot",
        "texts": [
            " (9) Finally, we get motion equation with one leg standing as: M(q)q\u0308 + h(q, q\u0307) + g(q) + Dq\u0307 = \u03c4 , (10) where, M(q) is inertia matrix, h(q, q\u0307) and g(q) are vectors which indicate Coriolis force, centrifugal force and gravity, D = diag[d1, d2, \u00b7 \u00b7 \u00b7 , d18] is matrix which means coefficients of joints\u2019 viscous friction and \u03c4 is input torque. If supporting-foot is surface-contacting and assumed to be without slipping, joint angle can be thought as q = [q2, q3, \u00b7 \u00b7 \u00b7 , q18]T . This walking pattern is depicted in Fig. 2 (a). When heel of supporting-foot should detach from the ground before floating-foot contacts to the ground as shown in Fig. 2 (b), the state variable for the foot\u2019s angle q1 be added to q, thus q = [q1, q2, \u00b7 \u00b7 \u00b7 , q18]T . Giving floating-foot contacts with a ground, contactingfoot like Fig. 3 appears with contacting-foot\u2019s position zh or angle qe to the ground being constrained. When constraints of foot\u2019s position and also foot\u2019s rotation are defined as C1 and C2 respectively, these constraints are represented by Eq. (11), where r(q) means the contacting-foot\u2019s heel or toe position in \u03a3W . C(r(q)) = [ C1(r(q)) C2(r(q)) ] = 0 (11) Then, robot\u2019s equation of motion with external force fn, friction force ft and external torque \u03c4n corresponding to C1 and C2 can be derived as: M(q)q\u0308 + h(q, q\u0307) + g(q) + Dq\u0307 = \u03c4 + jT c fn \u2212 jT t ft + jT r \u03c4n, (12) where jc, jt and jr are defined as: jT c = \u201e \u2202C1 \u2202qT \u00abT\u201e 1/ \u201a\u201a\u201a\u201a \u2202C1 \u2202rT \u201a\u201a\u201a\u201a \u00ab , jT t = \u201e \u2202r \u2202qT \u00abT r\u0307 \u2016r\u0307\u2016 , (13) jT r = \u201e \u2202C2 \u2202qT \u00abT\u201e 1/ \u201a\u201a\u201a\u201a \u2202C2 \u2202qT \u201a\u201a\u201a\u201a \u00ab "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure7-1.png",
        "caption": "Fig. 7. Rack-cutter.",
        "texts": [
            " When the rack-cutter fillet is positioned between both curves \u03b7 and \u03be (the arc AF2 of a radius \u03c12>\u03c11) an undercutting\u2014 type IIa appears, and when the rack-cutter fillet is placed between the curve \u03be and the line OE (the arc AF3 of a radius \u03c13>\u03c12) besides an undercutting \u2014 type IIa, an undercutting \u2014 type IIb is derived. The areas in which the rack-cutter fillet AFE (representing a circle, ellipse, trochoid, parabola, etc.), can be inscribed, without provoking an undercutting of the involute teeth of type IIa and type IIb, are defined by the boundary areas ACE and ADE, shown in Fig. 7. In this case, the boundaries AC and CE of the area ACE, constructed by Eq. (4), limit the undercutting of type IIa, and the boundaries AD and DE of the area ADE, constructed by Eq. (10), limit the undercutting of type IIb. The length of the common boundary AE of both boundary areas is equal to the tooth thickness sg over the tip straight line g-g of the rack-cutter. When on the middle-line m-m the tooth thickness sm is equal to the width em of the tooth space (sm=em=p/2), sg is calculated by the formula sg \u00bc m 0:5\u03c0\u22122h a tan\u03b1 ; \u00f014\u00de e depths h\u03b7 and h\u03be of the respective boundary areas ACE and ADE are derived by the parametric Eqs. (4) and (10) of the and th curves \u03b7 and \u03be, taking into consideration the equations X\u03b7 \u00bc X\u03be \u00bc r \u03b1 \u2212 cos\u03b1 sin\u03b1\u00f0 \u00de \u00fe 0:5sg ; \u00f015\u00de h\u03b7 \u00bc \u2212Y\u03b7 ; h\u03be \u00bc \u2212Y\u03be : \u00f016\u00de On Fig. 8 is shown the variation of the geometric shape of the boundary areas of the fillet curve (the areas ACE and ADE on Fig. 7) at different values of the pressure angle \u03b1 of the rack-cutter and different number of teeth z of the gear. It is immediately seen that by increasing the number of teeth z, the depths h\u03b7 and h\u03be of the respective boundary areas increase as well. The same depths, with one and the same number of teeth, decrease when increasing the pressure angle \u03b1. In the case where the rack-cutter fillet is profiled on an arc from a circle, the following boundary condition is defined: the undercutting \u2014 type II (IIa and IIb) is avoided if the radius of the fillet is smaller or equal to the radius of the curve \u03c1\u03b7,A (Fig. 7) of the boundary fillet \u2014 type IIa (curve \u03b7) in point A. Since in point A of the curve \u03b7 (Fig. 6) the value of the angular parameter is \u03c6=\u03b1, for the radius of the curve \u03c1\u03b7,A in this point according to Eq. (9), it is obtained \u03c1\u03b7;A \u00bc mz sin\u03b1 : \u00f017\u00de Then the boundary condition for a non-undercutting of type II, when the fillet curve of the rack-cutter is an arc \u043ef a circle of a radius \u03c1, is defined by the inequality \u03c1\u2264\u03c1max \u00bc \u03c1\u03b7;A \u00f018\u00de finally written as the following type(19) and is \u00f019\u00de where \u03c1*=\u03c1/m is a coefficient of the radius of the circle, on which the rack-cutter fillet is profiled",
            " It is seen from the figure that the teeth are not undercut of type I, as the tip line g\u2212g at x=xmin passes through point A'. Besides, the starting point b of the involute profile ba lies on the base circle, and as a result, the teeth are not undercut in a radial direction (\u03b4r=0 mm, \u03bbr=0%). In this case only a tangential undercutting is obtained, caused by the rack-cutter fillet AF, where \u03b4t=0.47 mm and \u03bbt=4.82%. This is determined by the fact that the rack-cutter fillet (profiled over a circle of a radius \u03c1*>\u03c1max\u204e=2.05) is placed in the area ADE (Fig. 7), outside the boundary area ACE. When generating the teeth, shown on Fig. 12, the parameters of the gear and rack-cutter, excluding the coefficient \u03c1*=10, are the same as on Fig. 11. And in this case, from Fig. 12 it is seen that the teeth are not undercut of type I, as the tip line g\u2212g passes through point A'. The undercutting obtained is of type IIb and is also provoked by the rack-cutter fillet AF. In this case, due to the larger value of \u03c1*, the fillet curve AF is positioned outside both boundary areas ACE and ADE"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000135_icra.2012.6225025-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000135_icra.2012.6225025-Figure8-1.png",
        "caption": "Fig. 8. Arm configurations tested for collision with a tea pot and a bunny",
        "texts": [
            " In contrast, our CD-ECoM algorithm uses the exact arm model of a continuum manipulator directly to avoid the high time and space cost of building and re-building mesh models for different configurations and thus facilitates motion planning for a continuum manipulator. Table II and Table III present the results of collision detection between the OctArm and two different object meshes, a tea pot mesh model with 1, 024 triangles and a bunny mesh model with 3, 851 triangles, at different OctArm configurations, as shown in Fig. 8. In configuration 1 and configuration 4, the arm is not in collision, and this is the most expensive case for collision detection with our CD-ECoM algorithm because no bounding volume of the object is used. On the other hand, OPCODE uses bounding volume hierarchies to speed up computation. As the result, the CD-ECoM algorithm takes more time than OPCODE for the coarse mesh 1 model of the arm. However, it is easy to use a bounding volume hierarchy for the object in our CD-ECoM algorithm to speed up the algorithm further"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003363_1.g002683-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003363_1.g002683-Figure1-1.png",
        "caption": "Fig. 1 The rotor hubwith pin joints for flap hinges and skewed lag-pitch hinges.",
        "texts": [
            " Section V describes test stand experiments and provides a comparison of measured motor torque, hub speed, blade lag, and blade flap cyclic responses to model predictions. Closing remarks in Sec. VII discuss how this work might impact the design and performance of future small UAV systems. The objective of the articulated rotor hub design is to allow modulating the drive shaft torque to induce controllable cyclic pitch variations. Our method is to kinematically induce a lag-pitch coupling through the combination of a conventional flap hinge and a skewed lag-pitch hinge. Figure 1 illustrates the physical device consisting of the hub, cross, blade grip, and blade bodies. The hub is attached to the cross by a flap hinge pin joint. The cross connects to the blade grip by a skewed lag hinge pin joint, and it is this skew angle that controls the degree of lag-pitch coupling. Similar hinge kinematics are depicted by Bousman [11] in the study of dynamic blade stability; now we exploit this structure as part of the control effector design. To this end the hub design is antisymmetric with a positive lag-pitch coupling imposed on one blade and a negative coupling imposed on the opposite blade. As a consequence, when a driving torque excites synchronous lead-lag motions in the two blades the pitch responses will be 180\u00b0 out of phase with each other. The addition of an explicit flap hinge is an improvement over previous work [9,10] at the expense of a small addition in design complexity. It is apparent from Fig. 1 that in the absence of the flap hinge the blade tip on the right would be forced to flap down both as the blade leads forward and lags backward about the skewed axis. At the same time, the blade tip on the left would be forced to flap upward twice per revolution as that blade obtains its maximum lead and lag angles. The resulting flapmotions for the two blades could not bematched, and these undesirable higher harmonics in the flapping response would contribute to large bending moments at the blade roots and unwanted airframe vibrations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure22.5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure22.5-1.png",
        "caption": "Fig. 22.5 Cylindrical model of the wound solid rotor (a) simplified scheme of the wound solid rotor cross-section; (b) wound solid rotor as a three-layer structure",
        "texts": [
            "30) allows the rotor slot wedge region to represent as a homogeneous conducting layer. 870 22 The Wound Solid Rotor Circuit Loops: Weak Skin Effect In magnetic terms, the rotor slot wedge region is represented as a magnetoanisotropic layer, of which the radial and tangential components of magnetic permeability can be determined by the expressions obtained in (16.25) and (16.27). Considering that the rotor joke region is represented as an isotropic conducting layer, the model of the wound solid rotor takes the form in Fig. 22.5b. The model represents a four-layer system, which includes the air gap layer with length equal to \u03b4/2, and the rotor layers reflecting the slot wedge, wound part of the tooth, and rotor joke regions. In this model, the wound part of the rotor tooth region is considered the active layer. Below, using the model shown in Fig. 22.5b, we consider the circuit loops of the wound solid rotor at a weak skin effect. In this work, the circuit loops of the rotor arise from consideration of its equivalent circuits, which can be constructed using the model shown in Fig. 22.5b. For this purpose, it is first necessary to have the equivalent circuits representing the corresponding layers of the rotor model in Fig. 22.5b. The layers of the rotor model can be replaced by T- or L-circuits. The air gap layer with length equal to \u03b4/2 is replaced by the equivalent circuit in Fig. 5.5. The T-circuit of the rotor slot wedge layer acquires the form shown in Fig. 20.6. The impedance values of this equivalent circuit (Zkl0,Z\u03c4kl1 and Z\u03c4kl2) can be defined by the expressions obtained in Chap. 20. In the rotor model in Fig. 22.5b, the wound part of the rotor tooth region is considered the active layer. On the basis of the equivalent circuit shown in Fig. 7.9 (Chap. 7), the T-circuit of the wound part of the rotor tooth layer takes the form in Fig. 22.6. The impedance values of this equivalent circuit (Z\u03a0z0,Z\u03c4\u03a0z1 and Z\u03c4\u03a0z2) are determined by the expressions given in (7.95). On the basis of expressions (7.97), (7.105), (7.107), (7.108), and (7.111), the resistance (r 2/s) and impedance Z 20 used in the equivalent circuit in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000099_j.triboint.2015.01.017-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000099_j.triboint.2015.01.017-Figure5-1.png",
        "caption": "Fig. 5. Heat transfer in an EHL contact.",
        "texts": [
            " Again, compared to the uncoated case, when the coated surface (ball) is moving faster, friction exhibits a more important variation (decrease). The origins of these observations are investigated next by carefully examining the localized thermal behavior of the contact. Before moving to a thorough thermal analysis of coated TEHD contacts, first it is essential to understand the concepts of heat generation and heat removal in any EHD contact (coated or uncoated). The concepts of heat generation and removal in elastohydrodynamic lubricated contacts are illustrated in Fig. 5. Heat is generated in an EHL conjunction by two separate mechanisms: compression and shear. The former is associated with the lubricant compression towards the inlet region of the contact. In fact, when compressed, any fluid would generate heat. However, in an EHL contact the lubricant is compressed on the inlet side of the contact (left side of Fig. 5), but then as it heads towards the outlet (right side) it is expanded. Therefore, a cooling effect is observed towards the outlet of the contact. This heating/cooling mechanism due to compression/ expansion generally has a negligible effect especially compared to shear heating. In fact, compression heating/cooling is only significant under pure-rolling conditions in the absence of any significant shear effects. Under rolling\u2013sliding conditions, heat is also generated in the lubricant film by shear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001706_s11071-018-4696-x-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001706_s11071-018-4696-x-Figure2-1.png",
        "caption": "Fig. 2 a Saturation nonlinearity and b dead-zone nonlinearity",
        "texts": [
            " x\u0302(k) indicates the state value saved by ZOH when the last event occurred. v\u0302(k) is the updated value of the controller when the last event occurred. v(k) is actual desired control effect. In order to transform the saturation nonlinearity into dead-zone nonlinearity, a dead-zone function is defined by sat(v(k)) = \u03c6(v(k)) + v(k), (9) with \u03c6(v(k)) = \u23a7\u23aa\u23a8 \u23aa\u23a9 \u2212 u0 \u2212 v(k) for v(k) < \u2212u0 0 for \u2212 u0 \u2264 v(k) \u2264 u0 u0 \u2212 v(k) for u0 < v(k). The saturation nonlinearity characteristics for sat(v(k)) and for \u03c6(v(k)) are shown in Fig. 2. The system (1) with the substitution of (8) and (9) is equivalent to x(k + 1) = Ai x(k) + Bi e(k) + Bi\u03c6(v(k)) + ai , (10) where Ai = Ai + Bi K i , ai = Bimi + ai . The following are definitions and lemmas used in the paper. Definition 1 Given a symmetricmatrix P i > 0, i \u2208 \u2118 and a scalar \u03c1 > 0, the set \u03be(P i , \u03c1) represents the ellipsoid \u03be(P i , \u03c1) = {x \u2208 R n : xT P i x \u2264 \u03c1}. (11) Definition 2 A region \u03c7 i is a region of asymptotic stability (RAS) with respect to the origin of system (1) if 0 \u2208 \u03c7 i and x(k) \u2192 0 as k \u2192 \u221e \u2200x0 \u2208 \u03c7 i "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001551_s106836661804013x-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001551_s106836661804013x-Figure6-1.png",
        "caption": "Fig. 6. Dependence of the coefficient of friction on the bearing temperature: (1) Ametist; (2) NIKA; (3) Izumrud; (4) T-900 fluoride solid lubricant coating [11].",
        "texts": [
            " The results of the evaluation of the relative life of lubricating materials using a spiral tribometer showed that cyclopentane greases (SH 050, SH 051, Rheolube 2000) have a longer life than greases based on PFPEs (PF 100, PF 101, Krytox 240AC) (Fig. 4). It is also noteworthy that Rheolube 2000 grease contained additives based on phosphates and phenols. It is noted that the use of molybdenum disulfide as an additive prolongs the life of greases based on cyclopentane by about 50%. Such an effect was not observed in the case of greases based on PFPEs, which is explained by the chemical reaction of molybdenum disulfide with components of PFPEs and the degradation of its lubricating properties [3]. In Fig. 6, the dependences of the coefficient of friction on the temperature for a radial ball bearing no. 6- YuT with different types of Russian greases based on PFPEs are presented. For comparison, the results of the measurements of the coefficient of friction of the L OF FRICTION AND WEAR Vol. 39 No. 4 2018 same bearing with a solid lubricant coating on the surfaces of holders are given. The tests were performed in the following conditions: the bearing dimensions were 9 \u00d7 26 \u00d7 8; the load was 1.02 kN; the type of motion was reciprocating rotational at the frequency of 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003242_ecmr.2017.8098681-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003242_ecmr.2017.8098681-Figure3-1.png",
        "caption": "Fig. 3. Representation of the main reference systems (colored triads) and transformations (dot lines). Our goal is the estimation of the Base-Pattern (BP ) and the Hand-Eye transformation (HE).",
        "texts": [
            " This proved to be useful to simplify the calibration process, which has to be done before every surgical procedure. Nevertheless, this approach is not as accurate as the common one. In [17], the approach is claimed to be accurate. Anyway, it uses the surgery instruments with known CAD models as calibration objects. In our scenario, an object with a known CAD model is less practical than a checkerboard. This section focuses on the description of our automatic procedure. For ease of understanding, the main reference systems and transformations are depicted in Figure 3. The symbols B, H , C and P are the initials of Base, Hand, Camera and Pattern, the names of the main reference systems. The symbols BH , HE, EP and BP stay for the respective rototranslations from the Base to the Hand, the Hand to the Eye, the Eye to the Pattern and the Base to the Pattern. Our goal is the estimation of BP and HE (in red). The automatic procedure can start after moving the robot to a starting position from which the calibration pattern is visible. Our calibration pattern is the asymmetrical circle pattern, currently supported in OpenCV [18], but could be also the classical black-white checkerboard with equally spaced squares"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001250_j.matpr.2017.07.234-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001250_j.matpr.2017.07.234-Figure2-1.png",
        "caption": "Figure 2: Capabilities of HLM",
        "texts": [],
        "surrounding_texts": [
            "Rapid Prototyping (RP) is a relatively old technology with the earliest references as old as 25 years ago. In 1988 the first commercial Rapid Prototyping system was born. The development was closely related to the development of applications of computers in industry. RP was a follow up of the development in CAD, CAM and CNC technologies. Rapid Manufacturing (RM), also known as Layered Manufacturing (LM), is a totally automatic process of manufacturing objects directly from their CAD models without the use of any tooling specific to the geometry of the objects being produced. RP adopts a divide-and-conquer approach in which the complex 3D object is split into several 2D slices that are simple to manufacture. Furthermore, as the object grows from bottom up, the chances of collisions are eliminated. In rapid manufacturing using deposition, the metal is deposited only in the required regions in a layer-by-layer manner. The material can be fed either in the form of wire or powder. The deposition technologies employ laser, electron beam or electric arc as the sources of thermal energy for melting the metal, in the order of their present popularity [1]. Table 01 lists various existing technologies in each category. Sajan Kapil / Materials Today: Proceedings 4 (2017) 8837\u20138847 8839"
        ]
    },
    {
        "image_filename": "designv11_13_0003114_lra.2017.2728200-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003114_lra.2017.2728200-Figure2-1.png",
        "caption": "Fig. 2: A quadruped robot labelled with the notation used",
        "texts": [
            " While promising, basic DRMs cannot be directly applied to grasping, manipulation or navigation of legged platforms, as these tasks require explicit contact with the environment and DRMs treat these contact points as collisions. To specifically address this issue, we introduce Contact DRMs (CDRMs) that in addition to maintaining workspace-configuration space mapping also map potential end-effector or foot-tip positions of the leg or manipulator arm. A legged robot R operates in a known workspaceW = R3, with the region O \u2286 W being occupied by obstacles. An example quadruped robot, along with the notation used, is shown in Fig. 2. The robot is composed of a robot body B to which n legs are attached. The body B is free to translate and rotate in SE(3). A full-body robot configuration is denoted by qr with the robot\u2019s configuration space (Cspace) C being the set of all possible configurations. The Cspace region invalidated by the presence of obstacles and selfcollisions is denoted by Cobs and the obstacle-free region of the configuration space Cfree = C \u2212 Cobs. For this paper, we assume that the n legs have the same kinematic structure, but this is not a hard limitation of the approach",
            " If this condition is required to be explicitly validated the path can be interpolated and static stability checked along the entire path. Due to proximity it is also assumed that the legs are collision free for body translation from Bnew to Bprev. Finally, any footholds that are not used as an element of a valid full-body state are discarded, as they are in inaccessible regions. This prevents subsequently generated states from attempting to connect to these footholds. The planner was tested using the quadruped shown in Fig. 2. Each of the four legs is kinematically identical, approximately 0.45m long and has three DOFs. Tests were run using a Core i7 4700M CPU and 16GB RAM. We first detail the generation of the CDRM in Sec. V-A, including the key parameters chosen. Planning results are then shown on planar terrains, individual challenging features and extended paths in complex scenarios which combine multiple of these features. We then compare our planner to a basic RRT implementation and a state of the art full-body legged motion planner",
            " These scenarios show the ability to plan extended motion paths which traverse a series of complex features. Most of the planning time is spent evaluating footholds, demonstrating the benefits of using the CDRM to optimise this. To validate our approach, we compared our CDRM planner with two other full-body planners (a) a naive 18-DOF (12 DOF leg configuration + 6 DOF body pose) RRT planner, and (b) the state of the art Reachability Based PRM (RBPRM) planner [4]. All three approaches were run on the robot kinematic model shown in Fig. 2. We evaluated the RRT planner on a planar terrain where samples that did not place all footholds on the ground were rejected. This took approximately 180s to generate a single valid sample, illustrating this approach is not feasible for planning with contacts as the probability of selecting valid contact configurations is small. 2377-3766 (c) 2017 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002465_amm.816.54-Figure16-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002465_amm.816.54-Figure16-1.png",
        "caption": "Fig. 16. The values and spots of the concentration of contact pressures for the angular position 75 o",
        "texts": [
            " Figure 13 presents the state of the Mises stresses with the enlarged view of the place of the occurrence of maximum stresses. The values of the reaction forces and their components are also presented. The values and spots of the concentration of contact pressures are presented on Figure 14. The position 75 o . Figure 15 presents the state of the Mises stresses with the enlarged view of the place of the occurrence of maximum stresses. The values of the reaction forces and their components are also presented. The values and spots of the concentration of contact pressures are presented on Figure 16. Table 1 presents the list of component reactions and their resultant reactions. This table also shows the forces and moments of friction which are needed to disassemble the connection. Figure 17 illustartes the graphic representation of the Mises stresses and contact pressures in function of the angular position of the shaft. Vertical reaction Ryc [N] 0.3686 0,4563 0,4278 0,3803 0,333 0,3043 (Ryl + Ryr) [N] 0.7606 0.7593 0.7608 0.7606 0.76 0.7603 Resultant reaction Rl [N] 2.169 1.049 0.649 0.312 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure11-1.png",
        "caption": "Fig. 11. The change of the resistance force on the surface of the spherical body, moving in the water flow.",
        "texts": [],
        "surrounding_texts": [
            "The drag coefficient is calculated by a classical formula [7]: Cd \u00bc 2 Fd q A U2 ; \u00f05\u00de where Fd \u2013 the drag force, obtained from simulation results, q \u2013 water density, U \u2013 mean velocity of a water flow, m/s, A \u2013 reference area, m2."
        ]
    },
    {
        "image_filename": "designv11_13_0001192_s10010-017-0250-0-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001192_s10010-017-0250-0-Figure1-1.png",
        "caption": "Fig. 1 Classical spiroid drive",
        "texts": [
            " Wenn die Anzahl der umh\u00fcllenden konischen Schneckenf\u00e4den kleiner ist, geschieht die Kr\u00fcmmungsst\u00f6rung im Allgemeinen nicht auf beiden Seiten eines Zahnes. Der Vermeidungsmechanismus der Kr\u00fcmmungsst\u00f6rung ist, dass die Grenzlinie w\u00e4hrend des realen Schneidgitters nicht existiert, da die Grenzlinie innerhalb der Einheit der umh\u00fcllenden konischen Schnecke liegt und ihre konjugierte Linie au\u00dferhalb der Einheit der Schleifscheibe liegt. Das numerische Ergebnis zeigt, dass der Zeh auf der e-Flanke das gr\u00f6\u00dfte potentielle Risiko hat, der Unter\u00e4tzung unterworfen zu werden. K As reported in the literature, the classical spiroid drive, as shown in Fig. 1, was initially invented by Oliver E. Saari at the Illinois Tool Works (ITW) in the 1950s [1]. In a classical spiroid gearing, the worm is conical in shape and the mating member is a face-type gear. The helicoidal surface of a classical spiroid is generally finishturned by a lathe tool with a straight-lined blade. The conical worm gear can be generated by a conical hob on conventional equipment after Olivier\u2019s second principle. Here, the conical hob is hence required to be identical to the mating spiroid"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002695_1.4033926-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002695_1.4033926-Figure1-1.png",
        "caption": "Fig. 1 Schematic representation of SPIF geometries. (a) Backing plate with a square opening, (b) square backing plate with a circular martenstic band on the sheet, and (c) backing plate with a circular opening. Top cone: wall angle 25 deg, inner diameter 175 mm, depth 20 mm and bottom cone: wall angle 50 deg, inner diameter 89.2 mm, depth 25 mm.",
        "texts": [
            " To increase the laser absorption and prevent laser reflection, which might affect accurate temperature measurement, the surface of the sheet metal was coated with graphite. After the determination of appropriate process parameters for laser transformation hardening, two-angled cone parts have been formed under the following circumstances: (1) using a backing plate with a square hole of side length 195 mm, (2) using the same backing plate with a circular martensitic band with diameter of 185 mm, and (3) using a backing plate with a circumferential orifice of diameter 182 mm (see Fig. 1). The second objective is to study the effect of generation of laser-affected hardened bands in the workpiece geometry during the SPIF process. In the corresponding set of experiments, the geometrical accuracy of a two-angled pyramid was studied while five martensitic bands were generated on the top pyramid (see Fig. 2). Finally, through accuracy measurements of the parts, the impeding role of these bands in preventing unwanted plastic deformation of the part, mostly at a slope change region, was studied",
            " An advanced three-dimensional FE is required to study this mechanism [16,25]. The further validation of FE modeling requires an in-depth microstructural and hardness analysis of the laser treated zone, which is discussed in Sec. 5. For both laser scanning conditions, the tailored microstructure zone was sectioned using wire-electrical discharge machining, parallel (XY plane) and transverse (XZ plane) to the laser scanning direction. For the large and small laser spot sizes, the hardened bands were cut from z\u00bc 0 (see Fig. 1(b)) and z\u00bc 10 mm (see Fig. 2), respectively. The samples were cold mounted in an epoxy resin and mechanically polished up to fine polishing using 1200 grit silicon carbide paper. The final stage of polishing was performed using 1 lm diamond paste; the samples were etched with 2 pct Nital to reveal the microstructure and gold coated for SEM investigations. OM images reveal that the microstructural changes show the same pattern for both laser scanning conditions; the laser-affected zone has thus been subdivided into a BM zone, a HAZ, a conversion zone (CZ), and a HZ for both conditions (see Figs",
            " The applied toolpaths were directly generated on this CAD geometry and no toolpath compensation strategies were used. The accuracy results of two case studies are presented below. 6.1 Hardened Band Used as Substitute for a Backing Plate. In the first strategy, the possibility of replacing the backing plate with a circular hardened band on the two-angled cone geometry is studied. The accuracy of the parts with and without circular backing plate (see Figs. 1(a) and 1(c)) is compared with the accuracy of the parts without backing plates but with a circular hardening band (see Fig. 1(b)). The hardened band is generated using the laser power of 200 W, scanning speed of 600 mm/min, and spot size of 6 mm and the forming is done by the tool diameter of 10 mm and the step down of 0.5 mm. Figure 12 illustrates the geometric accuracy results. It has been observed that the sheet metal part formed without circular backing plate (see Fig. 1(c)) shows prominent unwanted bending in the areas which have a maximum distance to the backing plate (see arrows in Fig. 12(a) and curve A in Fig. 14), the maximum deviation is 2.39 6 0.06 mm. This is due to the absence of a backing plate in the direct vicinity of the upper contour edge of the part which supports the sheet during the forming process. In normal SPIF processing, this type of over-forming can usually be prevented by using an appropriate backing plate (see Fig. 12(c) and curve C in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003128_miltechs.2017.7988729-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003128_miltechs.2017.7988729-Figure6-1.png",
        "caption": "Figure 6. Coordinate systems at center of mass of bodies 1, 2, 3 and i",
        "texts": [
            " The shooter \u2013 weapon system includes nine rigid bodies: \u2022 body 1 \u2013 shooter shoulder of mass m1, \u2022 body 2 \u2013 weapon frame of mass m2, \u2022 body 3 \u2013 breech system of mass m3, \u2022 body 4 \u2013 upper right arm (APBP) of mass m4, \u2022 body 5 \u2013 lower right arm (BPCP) of mass m5, \u2022 body 6 \u2013 right hand (CPDP) of mass m6, \u2022 body 7 \u2013 upper left arm (ATBT) of mass m7, \u2022 body 8 \u2013 lower left arm (BTCT) of mass m8, \u2022 body 9 \u2013 left hand (CTDT ) of mass m9 and It is necessary to take into consideration the mass mi of weapon moving parts. In order to analyze the dynamics of the shooter \u2013 weapon system, Descartes coordinate systems have been established at the mass center of each bodies and the whole system as shown in Fig. 6 and Fig. 7, where: O0 represents the stationary coordinate system fixed on the ground, Ok accounts for the local coordinate system established at the center of mass of k-th rigid body, { }for = 1, 2,..., 9k . The coordinates of system rigid bodies are labeled as follows: 1q - longitudinal displacement of body 1 along X0-axis, 2q - angular displacement of body 2 about X1-axis, 3q - angular displacement of body 2 about Y1-axis, 4q - angular displacement of body 2 about Z1-axis, 5q - longitudinal displacement of body 3 along X2-axis, 6q - angular displacement of body 4 about X1-axis, 7q - angular displacement of body 4 about Y1-axis, 8q - angular displacement of body 4 about Z1-axis, 9q - angular displacement of body 5 about Z4-axis, 10q - angular displacement of body 7 about X1-axis, 11q - angular displacement of body 7 about Y1-axis, 12q - angular displacement of body 7 about Z1-axis, 13q - angular displacement of body 8 about Z7-axis, 14q - angular displacement of body 6 about X5-axis, 15q - angular displacement of body 6 about Y5-axis, 16q - angular displacement of body 6 about Z5-axis, 17q - longitudinal displacement of body 6 along X2-axis, 18q - angular displacement of body 9 about X8-axis, 19q - angular displacement of body 9 about Y8-axis, 20q - angular displacement of body 9 about Z8-axis, 21q - longitudinal displacement of body 9 along X2-axis, iq - longitudinal displacement of body i along Xi-axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001267_1045389x17742731-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001267_1045389x17742731-Figure1-1.png",
        "caption": "Figure 1. (a) Smart three-lobe hole-entry hybrid journal bearing system, (b) developed three-lobe hole-entry journal bearing system surface, and (c) developed textured three-lobe hole-entry journal bearing system surface.",
        "texts": [
            " The contribution of this article is to numerically study the three-lobe hole-entry hybrid journal bearing having spherical dimples lubricated with ER lubricant. In this work, the bearing characteristics parameters are shown as a function of applied electric field and Sommerfeld number. Furthermore, the combined influence of surface texturing and ER lubricant on the dynamic behavior of the bearing is numerically performed in terms of linear trajectories for better understanding on the stability of bearing. General layouts of textured/non-textured three-lobe hole-entry hybrid journal bearings are illustrated in Figure 1(a) to (c). The general layouts also incorporate a subassembly demonstrating practical application of electric field in above-mentioned bearings. A strong electric field across fluid film domain is applied using direct current (DC) power supply (0\u20131200 V). Furthermore, the local Reynolds number for the flow in the clearance between surfaces is measured to be 0.0036, it means the inertia forces are negligible in comparison to the viscous forces. Therefore, laminar flow condition is used in the present analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001447_j.optlastec.2018.05.031-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001447_j.optlastec.2018.05.031-Figure1-1.png",
        "caption": "Fig. 1. The scheme of samples were cut off.",
        "texts": [
            " Secondly, determine the influence of the laser welding modes (KHLW and HCLW) on the welded joints formation and its behavior after welding. Thirdly, determine the effect of heat treatment on the weldability and the weld strength. The laser machine equipped with Yb laser source was used for laser sintering; beam diameter was 100 \u00b5m at laser power 195 W and scan speed 900 mm/s for 40 lm layer. The SLM samples were produced as follows; first a 100 100 10 mm sample was produced using powder of PH1 stainless steel. Further, the resulting sample was cut by electroerosion machine (EDM) onto plates of about 1.3 mm thick (Fig. 1). After EDM cutting the samples roughness was better than that after AM, which makes it possible not to prepare samples for welding. The chemical composition of SLM PH1steel and CR 321 type stainless steel are presented in Table 1. The cut AM samples were exposed to heat treatment under the following regimes NHT (no heat treatment), HT1 (1050 C for 90 min followed by furnace cooling), HT2 (650 C for 90 min followed by cooling in still air), HT3 (500 C for 90 min followed by cooling in still air)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003778_1.4042636-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003778_1.4042636-Figure4-1.png",
        "caption": "Fig. 4 The 3D model of Hooke joint",
        "texts": [
            " The angular velocity of each driving limb in fAg is expressed as Ax i \u00bc An i l i Ax B Ar Si\u00feAV BO \u00bc 1 l i An\u0302 i An\u0302 i Ar\u0302 Si AV BO Ax B h iT (45) Ax3 \u00bc An3 l3 AxB ArS3\u00fe AVBO \u00bc 1 l3 An\u03023 An\u03023 Ar\u0302S3 AVBO AxB T (46) The angular acceleration of each driving limb in fAg is written as Ae i \u00bc 1 l i An\u0302 i An\u0302 i A r\u0302 Si Aa BO Ae B h iT An\u0302 i Ax\u0302 TB Ar\u0302 Si Ax B 2 l 2i An\u0302 i An\u0302 i Ar\u0302 Si AV BO Ax B h iT An Ti Ar\u0302 Si An i T h i AV BO Ax B h iT (47) Ae3 \u00bc 1 l3 An\u03023 An\u03023 Ar\u0302S3 AaBO AeB T An\u03023 Ax\u0302T B Ar\u0302S3 AxB n o 2 l2 3 An\u03023 An\u03023 Ar\u0302S3 AVBO AxB Tn AnT 3 Ar\u0302S3 An3 T h i AVBO AxB T (48) The line velocity of the center of mass ZUi of oscillating rod in fixed coordinate system {A} is expressed as AV ZUi\u00bc Ax i An i lUi (49) AVZU3\u00bc Ax3 An3lUi (50) The linear acceleration of the center of mass ZUi of oscillating rod in fixed coordinate system {A} is expressed as Aa ZUi\u00bc Ae i An i lUi\u00fe Ax i \u00f0Ax i An i \u00delUi (51) AaZU3\u00bc Ae3 An3lUi\u00fe Ax3 \u00f0Ax3 An3\u00delUi (52) In the same way, the line velocity of the center of mass ZSi of the telescopic rod in the fixed coordinate system is expressed as AV ZSi\u00bc An i _l i\u00feAx i An i \u00f0l i lSi\u00de (53) AVZS3\u00bc An3 _l3\u00fe Ax3 An3\u00f0l3 lSi\u00de (54) The linear acceleration of the center of mass point ZSi of the telescopic rod in the fixed coordinate system is expressed as Aa ZSi\u00bc An i \u20acli \u00fe Ae i An i \u00f0l i lSi\u00de\u00fe Ax i \u00f0Ax i An i \u00de\u00f0l i lSi\u00de \u00fe2\u00f0Ax i An i \u00de _l i (55) AaZS3\u00bc An3 \u20acl3\u00fe Ae3 An3\u00f0l3 lSi\u00de\u00fe Ax3 \u00f0Ax3 An3\u00de\u00f0l3 lSi\u00de \u00fe2\u00f0Ax3 An3\u00de _l3 (56) 4.2 Force Analysis of Hooke Joint. The structure of Hooke joint can bear a certain torque. The 3D model of Hooke joint is shown in Fig. 4. The coordinate system O\u2013XYZ is established at Hooke joint. X and Y represents the direction of the two axis of Hooke joint, so the torque on the Hooke joint can be expressed as (Fig. 4) DX FX \u00bc DY FY \u00bc MUi (57) Where DX, DY are the vectors between the two ends of the Hooke joint axis. FX is the force that the driving limb acting on the Journal of Computational and Nonlinear Dynamics APRIL 2019, Vol. 14 / 041010-5 Downloaded From: https://computationalnonlinear.asmedigitalcollection.asme.org on 02/16/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use Hooke joint, FY is the force that moving platform acting on the Hooke joint. MUi is the torque of driving limb that form moving platform in the Hooke joint coordinate system established in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001520_0954409718789531-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001520_0954409718789531-Figure1-1.png",
        "caption": "Figure 1. Schematic of the railway vehicle with 31 DOFs.",
        "texts": [
            " This feature of a junction permits one to create a bond graph model only from kinematical or dynamical relations, while satisfying the other automatically and are thus used by several authors as a modelling tool for investigating different aspects of dynamics of rail and road vehicles.2,19\u201326 This unique feature of bond graph motivated the author to use bond graph formulation for modelling the railway vehicle in this work. Bond graph modelling of a high-speed railway vehicle with 31 DOFs running on curved track A railway vehicle is a combination of a number of components and wheelsets joined together through suspensions in longitudinal, lateral and vertical directions. Figure 1 illustrates model of the railway vehicle. It consists of a car body, bogies and two identical wheelsets for each bogie frame. The car body, bogie frames as well as the wheel sets are assumed as rigid bodies and are characterized by their mass inertia characteristics. Both the primary and secondary suspensions are modelled as a spring-damper combination having movement in the longitudinal, lateral and vertical directions. Thus, all possible motions of the system components are included in the analysis leading to a system with a total of 31 DOFs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001310_etep.2507-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001310_etep.2507-Figure8-1.png",
        "caption": "FIGURE 8 Submersible motor bypass pump system",
        "texts": [
            " However, in the process of submersible motor submersible oil, the flow velocity of the drilling fluid on the casing will be slowed down due to the change of the downhole pressure, and then the heat dissipation of the motor casing will be worse. Hence, to ensure heat transfer efficiency of submersible motor and the reliability of the oil\u2010road system, the bypass pump is introduced to install at the entrance of the submersible motor, which can increase the flow velocity of the drilling fluid indirectly when the oil that is provided by the submersible motor decreases. The submersible motor bypass pump system is shown in the Figure 8. Obviously, the exergy destruction rate minimum can be used as an objective function, which can be combined with the motor optimization algorithm to optimize internal temperature distribution of motor.25 Taking motor\u2010related parameters as optimization variables, the cyclic iterative optimization algorithm is used to prove the validity of exergy destruction rate minimum to heat transfer of the motor in the paper. To simplify the calculation, the outer diameters of stator and length of stator are just chosen as optimization variables, and other relevant data remain unchanged"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001101_j.procs.2017.05.324-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001101_j.procs.2017.05.324-Figure1-1.png",
        "caption": "Fig. 1. Dynamic model of a planetary gear system.",
        "texts": [
            " In view of this, this paper proposes a three-dimensional finite element model to precisely calculate the mesh stiffness of gear pairs under both healthy and tooth-cracked conditions. By incorporating the time-varying mesh stiffness into a lateral-torsionalcoupled nonlinear dynamic model, the vibration signals of a planetary gear train in two conditions are predicted for the fault diagnosis in the early crack stage. A lateral-torsional-coupled dynamic model of a planetary gear system is shown in Fig.1. The system is composed of one sun gear \u2018s\u2019, one ring gear \u2018r\u2019, one carrier \u2018c\u2019 and n identical planet gears denoted as pn. Herein, oxy is defined as the general coordinate system rotating at the speed of the carrier c with x axis passing through the center of the 1st planet. Each component i (i=s,r,c,pn) has three degree of freedoms, i.e., two translational motions (xi, yi) and one torsional motion (ui). The supporting stiffness for each component is represented by a lumped virtual spring with constant stiffness kij (j=x,y,u)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000983_access.2017.2669209-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000983_access.2017.2669209-Figure3-1.png",
        "caption": "FIGURE 3. Torque balance for DLIPM in the sagittal plane.",
        "texts": [
            " rm =  rmxrmy rmz =  lm sin \u03b8m cos\u03b1m lm sin \u03b8m sin\u03b1m lm cos \u03b8m =  zc 2 tan \u03b8m cos\u03b1m zc 2 tan \u03b8m sin\u03b1m zc 2  (3) rM =  rMxrMy rMz  =  2rmx 2rmy 2rmz + lM  (4) where lm is the distance between m and the contact point. Since rmz is constant, r\u0307mz is zero. Then, the first order differential equation and second order differential equation of rM are as follows r\u0307M =  r\u0307Mxr\u0307My r\u0307Mz  =  2r\u0307mx 2r\u0307my 2r\u0307mz  =  2r\u0307mx 2r\u0307my 0  (5) r\u0308M =  r\u0308Mxr\u0308My r\u0308Mz  =  2r\u0308mx 2r\u0308my 0 . (6) In order to derive the relationship between ZMP and CoM, the torque balance is applied. The x-direction torque balance is demonstrated in Fig. 3 and (7). RMx and Rmx are x-coordinate of massM and massmwith respect to the world coordinate, respectively. Mg(RMx \u2212 px)+ mg(Rmx \u2212 px) = Mr\u0308Mx \u00b7 rMz + mr\u0308mx \u00b7 rmz (7) As shown in Fig. 3, RMx and Rmx can be expressed as follows RMx = px + rMx = px + 2rmx (8) VOLUME 5, 2017 2461 Rmx = px + rmx = 1 2 px + 1 2 RMx . (9) After substituting (3)-(6) and (9) into (7), the torque balance equations become as follows R\u0308Mx ( Mzc + 1 4 mzc +MlM ) = ( M + 1 2 m ) gRMx \u2212 ( M + 1 2 m ) gpx (10) where R\u0308Mx = r\u0308Mx . Defining \u03c9\u0303n = \u221a (M + m/2)g/(Mzc +MlM + mzc/4), (10) can be rewritten as follows R\u0308Mx = \u03c9\u03032 nRMx \u2212 \u03c9\u0303 2 npx . (11) Similarly, the y-direction torque balance can be derived in the same way, so the dynamics of mass M can be expressed as (12)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000413_s12206-011-1201-6-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000413_s12206-011-1201-6-Figure6-1.png",
        "caption": "Fig. 6. New station of bearing.",
        "texts": [
            " According to the authors\u2019 past research [18, 19], radial stiffness of tapered roller bearing is ( ) ( ) 0 0 0.13 0.130 1.13 lim 1lim ln ln ( )ln 1 ln 0.13 r r r c F F r r r r r r r r r r r r r FK F F F F F Fn mF m F F C F F n m m F C F \u03b4 \u03b4\u0394 \u2192 \u0394 \u2192 \u2212 \u2212\u0394 \u2192 \u2212 \u0394 = \u0394 = +\u0394 \u2212 \u23a1 \u23a4\u2212 +\u0394 + + +\u0394 + \u23a2 \u23a5\u0394 \u0394\u23a3 \u23a6 = + + + \u22c5 (8) and radial damping of tapered roller bearing is 1 2 1C = 1 1+ C C \u2032 . (9) There is a tilt angle between inner and outer raceways besides radial compression; the tilt angle changes state vectors transfer relation of roller bearing, as Fig. 6 shows. Bearing station transfer relation considering bearing geometry structure is ( ) , 2 , R L i i i R L i i i R L L i i r i i R L L i i t i i i i Y Y Y M M k Q Q k i C m Y \u03b8 \u03b8 \u03b8 \u03b8 \u03c9 \u03c9 \u23a7 = = \u23aa = =\u23aa \u23a8 = +\u23aa \u23aa = + + \u2212\u23a9 (10) where L iY , L i\u03b8 , LMi and L iV denote the transverse displacement, slope, bending moment and shearing force at left side of station i respectively, while R iY , R i\u03b8 , R iM and R iV de- note the same quantities at the right side of station i respectively. tk and rk denote the bearing radial and bending stiffness respectively, \u03c9 denotes the rotating speed of the shaft and C denotes the radial damping of the roller bearing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003616_j.triboint.2018.09.024-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003616_j.triboint.2018.09.024-Figure1-1.png",
        "caption": "Fig. 1. Experimental apparatus picture.",
        "texts": [
            " To calculate surface free energy based on the Kaelble and Uy theory [29], the contact angles of ion-exchanged water and diiodomethane (Wako 1st grade, Wako Pure Chemical Industries, Ltd.) with a contact angle meter (DMs-401, Kyowa Interface Science Co., Ltd., Japan). To discuss influences of R, E, and vc on Ar and ar, distributions of real contact and film thickness between the rubber hemispheres and glass prism (084.4L100-45DEG-6P-4SH3.5, SIGMAKOKI Co., Ltd., Japan) were quantified for values of vc from 0.10 to 1.00mm/s in increments of 0.1 mm/s, and based on a total reflection method and light interferometry by using the original experimental apparatus as shown in the Fig. 1 [19,28]. The load cell (TL201Ts, Trinity-Lab Inc., Japan) and the electric cylinder (EACM4D30AZAC, Oriental Motor Co., Ltd., Japan) are already mounted to investigate relationship between friction and dewetting behaviors in the future. Ion-exchanged water was used as the lubricant. And the water depth was set at about 2.5 mm. The normal force increased to 0.0981 N, as the rubber got close to the grass. The pixel format, pixel size, and frame rate of a charge-coupled device (CCD) camera (AT-030MCL, JAI Ltd"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure23-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure23-1.png",
        "caption": "Fig. 23. Contact stress distribution at single and double contact position.",
        "texts": [
            " The difference shows that the stiffness of the carrier on the output side is more rigid than elsewhere. This effect is opposite to that of the twist deformation of the sun gear. As a consequence, the unevenly distributed stress of the sun-planet tooth pair due to the twist of the sun gear is thus compensated by the twist of the carrier with the planet shafts. As the result in Fig. 20 shows, the compensated stress distribution is almost the same with Case A in this study case. The corresponding contact patterns are illustrated in Fig. 23 (a). The characteristics of the contact stress distribution and the corresponding contact pattern of the tooth pairs are analyzed for different contact position of the gear pairs, including in normal tooth contact and tip corner contact. Two various cases are considered in the analysis: the sun and the planet gear are either in single pair tooth contact or in double tooth contact. The saddle-shaped distributed contact stresses of the tooth pairs are simulated by the proposed approach, where tip corner contact does not occur, as illustrated in Fig. 23 . The variation of the stress along the major axis of the contact pattern is uneven in the annulus-planet tooth pairs, and more uniform in the sun-planet, as the similar results show in Figs. 17 and 18 . Because of the curvature relation of concave-convex contact, the minor axis length of the annulus-planet tooth pairs is larger than that of the sun-planet, but the contact stress is smaller. The contact stresses are analyzed for various contact positions before the beginning and after the end of normal contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000251_icar.2013.6766543-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000251_icar.2013.6766543-Figure1-1.png",
        "caption": "Fig. 1: Active-caster mechanism(a), schematic top view(b)",
        "texts": [
            " Together with the accurate self-positioning system for the omnidirectional robots, the proposed navigation method is verified by a series of experiments using the robot prototypes. II. OMNIDIRECTIONAL ROBOT WITH ACTIVE-CASTERS The authors group developed a unique omnidirectional drive system called active-caster. The drive system involves a drive wheel mechanism which forms a caster configuration in which a wheel axis and a steering axis are driven by electric motors to create omnidirectional motion by emulating caster behaviors. Figure 1(a) shows an overview of active-caster[5]. This is a single wheel drive mechanism in which a normal tire is off-centered from a center of the steering axis. The mechanism equips with two motors for actuating the wheel shaft and the steering shaft independently. As is seen in Fig.1 (b), when only the wheel shaft is rotated by the first motor, the caster moves in a forward direction. While only the steering shaft is rotated by the second motor, the mechanism rotates about the point of contact. By this rotational motion, the steering shaft moves in the tangential direction of the circle whose center is at the point of contact with a radius s (casteroffset). These velocity vectors are independently controlled and directing right angle to each other. Therefore, a resultant velocity V is generated at the center of the steering shaft which direction is determined by the ratio of two velocities, which can direct in an arbitrary direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000301_s12555-012-0243-6-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000301_s12555-012-0243-6-Figure1-1.png",
        "caption": "Fig. 1. Kinematic relationship of the CAES.",
        "texts": [
            " Development of the Cartesian Arm Exoskeleton System (CAES) using a 3-axis Force/Torque Sensor 977 of the CAES while maintaining a basic posture defined by the operator. In the external load lifting experiment, we showed that the operator can lift external loads with a small force. The experimental results demonstrated the effectiveness of the proposed methods. 2. THE CARTESIAN ARM EXOSKELETON SYSTEM (CAES) 2.1. The kinematic relationship of the CAES The Cartesian arm exoskeleton system (CAES) has five-degree-of-freedom (5-DOF), as shown in Fig. 1. There are three pitch joints, \u03b82, \u03b83, \u03b84, and two roll joints, \u03b81, \u03b85. The basic posture of the CAES is a bent posture. It maintains 0.3m along the X-axis, zero along Y-axis and - 0.3m along the Z-axis. Using this bent posture, we can secure the workspace and prevent a singularity. The final goal of the CAES is an exoskeleton system that may be worn on the whole body. The CAES is one of the test-beds, especially with a single arm. Therefore, the CAES has a similar configuration to the bent human arm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002312_1.4029054-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002312_1.4029054-Figure3-1.png",
        "caption": "Fig. 3 Ball bearing model",
        "texts": [
            " The rotor\u2019s rotational equation can be given: Ip \u20achr \u00bc nbx1Mb1 nbx2Mb2 (11) where Mbk denotes the frictional torque generated in ball bearing (k\u00bc 1, 2 mean the left and right sides). This section mainly describes the detailed mechanical model of the rotor and dynamic equations in 5-DOF. 3.2 Ball Bearing Model. Taking the ball bearing at left side of the rotor (shown in Fig. 2), for example, the motion of the inner race is consistent with that of the rotor at point B. In order to analyze the ball bearing\u2019s motion, its mechanical model is depicted (Fig. 3). In Fig. 3, Od, Oo, and Ob represent the centers of the ACABD, the ball bearing\u2019s outer race and the rotor at point B, respectively; _ho means the rotational speed of the outer race; xo and yo separately denote the x- and y-displacements of the outer race; xb and yb separately represent the x- and y-displacements of the rotor at point B, which can be transformed to the rotor displacement vector xb yb \u00bc 1 2lc lc \u00fe lb 0 lc lb 0 0 lc \u00fe lb 0 lc lb xs1 ys1 xs2 ys2 0 BB@ 1 CCA (12) To simplify calculation, we assume that there always exists a ball located at the normal direction of the impact force when each impact occurs [22]",
            " Tv means the viscous frictional torque of the ball bearing Tv \u00bc 160 10 7Nbf0d3 m tnb < 2000 10 7Nb f0 tnb\u00f0 \u00de2=3d3 m tnb 2000 ( (18) where f0 is the factor depending on the bearing type and the mode of lubrication; t represents the kinematic viscosity of the lubricant for the ball bearing at operation temperature; nb denotes the relative speed difference between the inner and outer races. This section calculates the internal impact force and frictional torque in the ball bearing. 3.3 Contact Model. As shown in Fig. 3, the clearances between each support and the outer race are unequal. Among these clearances, there must exist a smallest value. That means on each support, a particulate point keeps the closest distance to the outer race. Because of the ACABD\u2019s symmetric structure, the distances from these points to the outer race are consistent. During the impact process, the outer race would first contact with the supports at these nearest points. Connecting these points would achieve an equivalent clearance circle. From the working principles of ACABD, the distances from these points to the outer race would decrease with the swing of the supports, which means the radius of this equivalent clearance circle reduces at the same time. Therefore, the impact between the outer race and the supports can be simplified to that the outer race impacts with a variable-radius bearing housing hole. In the initial state of the ACABD, seen from Fig. 3, the circle is composed of the midpoint of each support\u2019s surface. We determine this circle as initial clearance circle, namely, dashed circle depicted in Fig. 3. The radius of this initial clearance circle is Rc0 \u00bc Ro \u00fe q0 (19) where Ro is the radius of the outer race; and q0 is the initial protective clearance. The ACABD is designed to reduce the radius of the equivalent clearance circle. In order to analyze the influences of different supports on the reduction of this clearance circle\u2019s radius, supports with concave and convex shape surfaces are studied, as shown in Fig. 4. Suppose the section of support\u2019s surface is an arc with the center of curvature Om and radius Rm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003887_j.ymssp.2019.04.039-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003887_j.ymssp.2019.04.039-Figure3-1.png",
        "caption": "Fig. 3. Lumped parameter model for planetary gears and parallel gears.",
        "texts": [
            " u3b and ug denote the respective rotational displacement of Stage 3 gear b and the generator at their respective base circle radius r3b and rg . Let Tg denote the external torque loaded to the generator rotor. Let ke4 denote the torsional stiffness of the spring that represents the shaft connecting Stage 3 gear b and the generator rotor. Both Stage 1 and Stage2 are planetary gear stages. Each of the two stages comprises of one carrier as input, one stationary ring gear, one sun gear as output and three equally spaced planet gears (cf. Fig. 3). The relative displacement between the carrier and the ith planet (i = 1,2,3) is represented by dncix \u00bc xni xnccosuni yncsinuni dnciy \u00bc yni \u00fe xncsinuni ynccosuni unc ( \u00f016\u00de where the subscript n \u00bc 1;2 denotes the stage index; the subscript f \u00bc c; r; s; i denotes the carrier, ring, sun, and the ith planet respectively. xnf ynf; \u00f0f \u00bc c; r; s\u00de refers to the transverse displacement of the specified carrier, ring gear and sun gear; xnf ynf; \u00f0f \u00bc i\u00de refers to the transverse displacement of the ith planet relative to the carrier",
            " The un-damped differential equations for the DOFs of Stage 1 are represented by m1c\u20acx1c \u00bc k1cxx1c \u00fe P3 i\u00bc1k1cixd1cixcosu1i P3 i\u00bc1k1ciyd1ciysinu1i m1c\u20acy1c \u00bc k1cyy1c \u00fe P3 i\u00bc1k1cixd1cixsinu1i \u00fe P3 i\u00bc1k1ciyd1ciysinu1i I1c \u20acu1c r21c \u00bc ke1 u1c r21c ut rt r1c \u00feP3 i\u00bc1k1ciyd1ciy 8>>>< >>>: \u00f025\u00de m1r\u20acx1r \u00bc k1rxx1r P3 i\u00bc1k1rid1risinu1ri m1r\u20acy1r \u00bc k1ryy1r \u00fe P3 i\u00bc1k1rid1ricosu1ri I1r \u20acu1r r21r \u00bc k1rtu1r \u00fe P3 i\u00bc1k1rid1ri 8>>>< >>>: \u00f026\u00de m1s\u20acx1s \u00bc k1sxx1s P3 i\u00bc1k1sid1sisinu1si m1s\u20acy1s \u00bc k1syy1s \u00fe P3 i\u00bc1k1sid1sicosu1si I1s \u20acu1s r21s \u00bc ke2 u1s r21s u2c r2c r1s \u00feP3 i\u00bc1k1sid1si 8>>>< >>>: \u00f027\u00de m1i\u20acx1i \u00bc k1cixd1cix k1sid1sisina1si \u00fe k1rid1risina1ri m1i\u20acy1i \u00bc k1ciyd1ciy k1sid1sicosa1si k1rid1risina1ri I1i \u20acu1i r2 1i \u00bc k1sid1si k1rid1ri 8>< >: i \u00bc 1;2;3\u00f0 \u00de \u00f028\u00de The un-damped differential equations for the DOFs of Stage 2 are represented by m2c\u20acx2c \u00bc k2cxx2c \u00fe P3 i\u00bc1k2cixd2cixcosu2i P3 i\u00bc1k2ciyd2ciysinu2i m2c\u20acy2c \u00bc k2cyy2c \u00fe P3 i\u00bc1k2cixd2ixsinu2i \u00fe P3 i\u00bc1k2ciyd2ciysinu2i I2c \u20acu2c r2 2c \u00bc ke2 u2c r2 2c u1s r2c r1s \u00feP3 i\u00bc1k2ciyd2ciy 8>>< >>: \u00f029\u00de m2r\u20acx2r \u00bc k2rxx2r P3 i\u00bc1k2rid2risinu2ri m2r\u20acy2r \u00bc k2ryy2r \u00fe P3 i\u00bc1k2id2ricosu2ri I2r \u20acu2r r22r \u00bc k2rtu2r \u00fe P3 i\u00bc1k2rid2ri 8>< >: \u00f030\u00de m2s\u20acx2s \u00bc k2sxx2s P3 i\u00bc1k2sid2sisinu2si m2s\u20acy2s \u00bc k2syy2s \u00fe P3 i\u00bc1k2sid2sicosu2si I2s \u20acu2s r2 2s \u00bc ke3 u2s r2 2s u2c r2c r2s \u00feP3 i\u00bc1k2sid2si 8>>< >>: \u00f031\u00de m2i\u20acx2i \u00bc k2cixd2cix k2sid2sisina2si \u00fe k2rid2risina2ri m2i\u20acy2i \u00bc k1ciyd2ciy k2sid2sicosa2si k2rid2risina2ri I2i \u20acu2i r2 2i \u00bc k2sid2si k2rid2ri 8>< >: \u00f0i \u00bc 1;2;3\u00de \u00f032\u00de According to the scheme of Stage 3 parallel gears (cf. Fig. 3), the relative displacement and the relative velocity over the line of action of gear a \u2013 gear b meshing are represented by d3ab \u00bc x3asinu3ab y3acosu3ab u3a \u00fe x3bsina3ab \u00fe y3bcosa3ab u3b _d3ab \u00bc _x3asinu3ab _y3acosu3ab u3a \u00fe _x3bsina3ab \u00fe _y3bcosa3ab _u3b ( \u00f033\u00de where the subscript n \u00bc 3 denotes the stage index. The subscripts f \u00bc a; b denote the drive gear and driven gear respectively; xnf ynf denotes the transverse displacement of the specified part. unf denotes the rotational displacement at its base radius rnf"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000281_1.3657244-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000281_1.3657244-Figure1-1.png",
        "caption": "Fig. 1 Toroid rollers",
        "texts": [
            " These were (1) a sufficiently large number of tests, (2) extreme precision in roller fabrication, (3) a rigidly maintained test procedure, (4) a high degree of lubricant cleanliness and temperature control, and (5) statistical preplanning and analysis. This report presents the results of an investigation which incorporated the afore-mentioned requirements. Method of Test Contact roller tests for antifriction bearing studies are performed with a toroid surface rolling on a cylindrical surface, as shown in Fig. 1. This conformity simulates the rolling action of a ball in a race, a wheel on a rail, or crowned gear teeth in pitch line mesh. The simple toroid arrangement provides rolling unaffected by retainers or neighboring dynamic elements. The contact stresses in toroid rolling are of particular interest, as these may lead to fatigue failure of the contacting surfaces. The theoretical definition of these stresses is somewhat inexact because of the complexity of the interrelation of load and elastic behavior of materials under dj'namic conditions",
            " tact under static normal load; while the formulas of Thomas and Hoersch derive the subsurface shear stresses, again under static conditions. Modern theoretical developments, such as those of Lundberg and Palmgren, reference [2]; Johnson, reference [3]; FesslerandOllerton, reference[4]; Moyar,reference[5],and Moyar and Morrow, reference [6], define the stresses generated by dynamic loading. Of these, the orthogonal shear-stress approach of Fessler and Ollerton appears most practical and was applied, in addition to the static definitions, to the results presented in this report. Contact rollers of the general shape shown in Fig. 1 were tested to failure at ambient room conditions. Three toroid radii, three maximum shear-stress levels, and three steel heats were selected. The stresses and contact areas were determined by the methods presented in references [7] and [4]. Both the maximum shear stress, which is a function of static load conditions, and the orthogonal shear-stress range, which is a function of the transverse movement of load in the contact surface, were computed. The constants pertinent to each toroid shape are shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000053_1464419314566086-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000053_1464419314566086-Figure5-1.png",
        "caption": "Figure 5. Static equilibrium of SRB.",
        "texts": [
            "comDownloaded from In the above equations, is the azimuth angle of No. 1 roller. The induced roller thrust loads in z direction are in the state of self-balance.33 Elastic deformation analysis of roller raceway with no roller error When the rollers are symmetrically distributed in the range of load zone (shown in Figure 4), the SRB inner ring will move along the loading direction under an applied radial load Fr. When the rollers are not symmetrically distributed in the range of load zone (shown in Figure 5), instead of moving along the loading direction, the inner ring will move to a direction that has an angle of \u2019 relative to the loading direction. The value of \u2019 can be determined by simultaneously solving the static equilibrium equations both in x direction and in y direction. According to the deformation compatibility relationship of roller raceway, the normal elastic deformation of roller raceway 1 j and 2 j can be expressed as 1 j \u00bc r cos\u00f0 1j \u00fe \u2019\u00de cos Pd 2 \u00f025\u00de 2 j \u00bc r cos\u00f0 2j \u00fe \u2019\u00de cos Pd 2 \u00f026\u00de where, 1j and 2j are the jth roller azimuth angle in the first and second row respectively 1j \u00bc 2 Z \u00f0 j 1\u00de \u00f027\u00de 2j \u00bc 0 \u00fe 2 Z \u00f0 j 1\u00de \u00f028\u00de Z is the number of rollers for each row, 0 is the staggering angle between No"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002889_s40436-016-0158-1-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002889_s40436-016-0158-1-Figure8-1.png",
        "caption": "Fig. 8 The vibration mode of the spindle unit under thermal steady state",
        "texts": [
            " With the simulation results plugging into Eqs. (17) and (18) and the diagram provided in Ref. [16], we can get the additional load of the front or the rear bearings that is generated by thermal deformation. Then the bearing stiffness under the thermal static state can be calculated by Eqs. (15) and (16). Since the maximum thermal deformation is about 0.06 mm which is relatively small in comparison with the size of the spindle unit, it is ignored when the spindle mode is analyzed. The numerical result is shown in Fig. 8, and the first order natural frequency of the spindle unit is 828 Hz and the second one is 1 940 Hz. Compared with the static state, the vibration modes are the same, and the relative deformation is not very large. It shows that the thermal state of the motorized spindle system mainly affects the natural frequency of the rotor. The numerical simulation step can be summarized in Fig. 9. It is important to consider many factors when choosing the experimental instruments. The purpose of the experiments is to investigate the thermal characteristic impact on the modal characteristic"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003697_j.ijthermalsci.2018.11.027-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003697_j.ijthermalsci.2018.11.027-Figure10-1.png",
        "caption": "Fig. 10. Mechanical spindle diagram in Workbench.",
        "texts": [
            " In order to improve the quality of meshing and to save simulation time, the following simplified treatments were implemented: 1) the screw hole, tubing, oil hole and other small features of the spindle were ignored; 2) the chamfer and fillet were removed; 3) the simple structure of solid model was adopted to replace component of complex structure, for example, using a simple ring instead of the spindle cooling jacket. The simplified mechanical spindle model was imported into ANSYS-Workbench, as shown in Fig. 10. The material properties of bearing-spindle system components were set according to the spindle material 38CrNoALA, the bearing material GCr15 and the remaining parts of 45 steel. At rotation speed of 10000r/min, 84.50W heat source was applied to the front and rear bearings of the bearing-spindle system. For convenience, in this paper the direct loading method was used and 42.25W heat source was applied to bearing's inner and outer raceway in the form of heat flux density. From the simulation results, the maximum temperature of the bearing-spindle system is 43",
            "68 (10) Based on the optimized thermal coefficients by multi-objective optimization, the corresponding thermal convection parameters are substituted into the bearing-spindle thermal elongation model to obtain the modified thermal elongation in the axial directions. As shown in Fig. 11(b), the maximum thermal elongation of the spindle's right and left end face (X direction) are 12.59 \u03bcm and 9.99 \u03bcm respectively, and the total thermal elongation in the axial direction is: \u2032 = + = + =+ \u2212L L L \u03bcm\u0394 12.59 9.99 22.58 (11) 5.2. Experimental verification of bearing-spindle thermal elongation The test bench of high-precision bearing-spindle system was set up, whose structure is the same as shown in Fig. 10. The high precision laser displacement sensor was fixed on the surface of the test bed with the magnetic table. The test precision is 0.1 \u03bcm and the axial thermal elongation of the spindle was tested by the displacement. Since the displacement sensor was fixed on the surface of the test bed, the measured thermal elongation is the absolute thermal elongation of the spindle (as in Fig. 12). Fig. 13 shows the spindle thermal elongation data at 10000 r/min within 3 h. The black line represents the original signal and the red is the trend curve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001049_1.4036586-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001049_1.4036586-Figure1-1.png",
        "caption": "Fig. 1 The dimensions of the CT specimen",
        "texts": [
            "org/about-asme/terms-of-use surfaces of the repaired and unrepaired specimens at the crack tip were examined by scanning electron microscopy (SEM) equipped with an energy dispersive spectrometer (EDS). Microstructure and fracture behavior were analyzed and reported. 2.1 Material and Specimen Preparation. The chemical composition of the precracked 304SS compact tension specimens used in the experiment is listed in Table 1. The mechanical properties of 304SS are presented in Table 2. The CT specimens were prepared in accordance with ASTM E647-12 standards. The dimension of the specimen with a thickness of 5 mm is shown in Fig. 1. A computer-controlled cutting machine was utilized to create a notch in the specimen. The precrack was created by a wire electric discharge machining system, with a length of 2 mm and a width of 0.2 mm. In this study, a mixture of commercially available micron-sized 304SS powder and nano-WC powder was used as feedstock material in the laser repair. Powder ingredients of 304SS are selected similar to those of the specimen, as listed in Table 1. The sizes of 304SS particles are within the range of 30\u201350 lm, and the sizes of nano-WC particles are around 50 nm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002411_ssd.2015.7348200-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002411_ssd.2015.7348200-Figure3-1.png",
        "caption": "Fig. 3. Implemented plug/socket.",
        "texts": [
            " RSI allows the realization of sensor guided robot motions. It is an additional module which realizes real time signal processing and the access to the position control loops. It is possible to create relative complex controller structures. After having created the whole RSI structure, it is able to run in real time with the interpolation cycle in parallel with the standard KRL (KUKA Robot Language) program. The proposed algorithm has been tested on plug/socket type IEC 60309 which is very hard for plugging even for the human, as shown in Fig. 3. IEC 60309 plug/socket has many safety features which make plugging task almost very difficult. It can be seen in Fig. 3, the IEC 60309 industrial plug/socket is very secure and weatherproof, because the plug should cover the whole socket cavity. Furthermore, pins and slots dimensions are almost the same to permit only proper insertion of plug into socket and the ground pin has a larger diameter than the other pins, preventing the wrong type of plug being inserted in a socket. In addition to that, the plug could not be inserted inside the socket unless the major keyway on the plug aligns with the notch on the socket at the beginning"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure10-1.png",
        "caption": "Fig. 10. Undercutting \u2014 type I: z=6; x=\u22120.2bxmin; \u03b4r\u204e=0.125; \u03b4t\u204e=0.147; \u03b1=20\u00b0; ha\u204e=1; c*=0.25; \u03c1*=0.38.",
        "texts": [
            " In the third column of the same table the formulas for the calculation of the respective values are provided. The generation of the involute profile and fillet curve of the cut teeth is realized with the help of the approach used by Litvin [1], Fetvaci [21] and other authors, under which the relative positions of the rack-cutter are constructed (drawn) in the plane of the gear being cut. For this purpose a software product was developed by the authors. On Figs. 10, 11 and 12 the tooth generation of the three gears in Table 2 is visualized. From Fig. 10 it is immediately seen that if the rack-cutter parameters are \u03b1=20\u00b0, ha\u204e=1, c*=0.25, \u03c1*=0.38, when cutting a gear of parameters z=6 and x=\u22120.2 an undercutting of type I is obtained, because the tip line g\u2212g of the rectilinear profile of the rack-cutter does not cross the meshing line in the part A'PK. In this case, the teeth are undercut simultaneously in a radial and tangential direction and the undercutting indices are respectively: \u03b4r=1.25 mm; \u03bbr=12.74%; \u03b4t=1.47 mm; \u03bbt=20.66%. The presence of anundercutting of type II\u0430 is found from Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002939_access.2017.2652984-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002939_access.2017.2652984-Figure2-1.png",
        "caption": "Fig. 2 Collision of MAF",
        "texts": [
            " The condition under which the formation has formed the target configuration is defined as ,i i cic q i     (3) According to (1), (2), and (3), we can get 2 22 2 2 , , ,i j ij i i j j c ijc q c q c q k i j i j          (4) Supposing 1c c cnk k k   , ck denotes the formation configuration coefficient. The smaller ck , the more rigid the formation configuration. According to (4), we can get the condition under which the formation has formed the relative target configuration as follows. , , ,i j ij c ijc q k i j i j      (5) Definition (3). The \u201ccollision\u201d in this paper means that the distance between i and j is less than or equal to the relative safe distance during the formation configuration forming process of MAF (as shown in Fig. 2), namely : , , , ,i j ij sij i j ij c ijFC d d c q k i j i j       (6) where the collision is denoted by i jFC . It should be noted that the \u201ccollision\u201d in this paper means the potential risk of collision, instead of meaning 0ijd  . 1) The members of MAF can obtain the information of local absolute positions and velocities through the navigation system in real time. 2) The communication network of MAF is fully-connected through the one hop or multi-hop pattern. As shown in Fig. 2, since the relative target configuration position ijc is approximately opposite to the relative position ijq , the distance ijd will be quickly reduced by the formation configuration control algorithm. When ij sijd d , ijd will 2169-3536 (c) 2016 IEEE. Translations and content mining are permitted for academic research only. Personal use is also permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003379_0954410018764472-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003379_0954410018764472-Figure5-1.png",
        "caption": "Figure 5. ANSYS mesh. Figure 6. Morphing process.",
        "texts": [
            ", 0:8 as the reference points, then the corresponding aerodynamic parameters for these configurations can be approximately calculated through computational fluid dynamics (CFD). Firstly, the outline dimensions of the aircraft are obtained by observation. Then the three-dimensional model of morphing aircraft under different configurations is established by Catia. After gridding model and importing it into fluent software of ANSYS, the aircraft aerodynamic force and moment can be calculated (shown in Figure 5). Finally, we integrate, analyze, and calculate the obtained data of those static configurations and the expressions of aerodynamic derivatives of the morphing aircraft during the whole morphing process can be derived by least square fitting of Matlab. Therefore, we get the continuous aerodynamic parameter expressions in Appendix 5. Next, we take the following morphing process (shown in Figure 6) as an example to illustrate the relationship between the aerodynamic coefficients and aerodynamic shape"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002312_1.4029054-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002312_1.4029054-Figure14-1.png",
        "caption": "Fig. 14 Three tested ACABDs",
        "texts": [
            " Journal of Engineering for Gas Turbines and Power JUNE 2015, Vol. 137 / 062502-9 Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use The structural parameters are consistent with the simulation parameters shown in Tables 1\u20133. The initial radial clearance between ACABD and the outer race of the ball bearing is 0.15 mm. And the radial clearance between radial magnetic bearing and the rotor is 0.4 mm. The three tested ACABDs are shown in Fig. 14. The experimental facilities are shown in Fig. 15. In these experiments, when a 5DOF AMB system has operated normally to suspend the rotor stably, we cut the power supply of the singleside electromagnet to realize the rotor\u2019s unilateral dropping. The rotor and outer race\u2019s speeds are detected by two fiber sensors separately. The displacement signal of the rotor is collected by the data acquisition card. The frequency converter is used to control the rotational speed of the motor in the AMB. 5.2 Test Results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000952_iet-epa.2016.0617-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000952_iet-epa.2016.0617-Figure8-1.png",
        "caption": "Fig. 8 Experimental setup specifications",
        "texts": [
            " Therefore at high load\u2013torques, efficiencies of the proposed method and MTPA method are closer to each other (although, the proposed method has more efficiency). According to (24), optimum value of id for loss-minimisation depends on motor parameters including lf , Ld, Lq, Rs and Rc. Different losses against id are shown in Fig. 7 in speed of 2000 rpm and load of 8.5 N m. A laboratory setup is constructed in order to verify the proposed method applicability. Test setup configuration and experimental setup are illustrated in Fig. 8. Interior permanent-magnet synchronous motor (IPMSM) parameters are listed in Table 1. IET Electr. Power Appl., 2017, Vol. 11, Iss. 3, pp. 447\u2013459 & The Institution of Engineering and Technology 2017 Fig. 9 Fe loss resistance and mechanical and stray losses measurement (experimental result) a Semi-input power against square of speed emf in 2000 rpm \u2013 v2 e l2d + l2q ( ) is in power invariant transformation b Semi-input power against square of speed emf in 750 rpm \u2013 v2 e l2d + l2q ( ) is in power invariant transformation c Sum of mechanical and stray components of torque (Tf) in different speeds dc-Link voltage and switching frequency are equal to 250 V and 15 kHz, respectively. A microchip dsPIC microcontroller with part number of dsPIC30f4011 is utilised in experimental setup. As shown in Fig. 8b, IPMSM is coupled to dc-generator as load. Results are sampled via an Advantech USB-4711A data acquisition apparatus. Also, a \u2018Digital Harmonics Monitor: Model HWT-2000\u2019 device is utilised at the input of diode\u2013rectifier to measure input power of whole drive. Range of analysed harmonic orders is from fundamental to the 40th harmonic, and its resolution is 14 bit. Also measuring current and voltage ranges can be adjusted on 5\u2013300 A and 150\u2013500 V, respectively. Also, in order to compare the proposed method with another loss-minimisation method, optimum values of iod (as a function of speed and iq) are calculated and are put into look-up tables; therefore, in order to verify negligible impact of utilised approximations, the proposed method is compared with an offline look-up table-based method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000600_0954406213477777-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000600_0954406213477777-Figure4-1.png",
        "caption": "Figure 4. The gearbox and its assembly.",
        "texts": [
            "comDownloaded from Experiments were performed on a Machinery Fault SimulatorTM (MFS) and a schematic diagram of it is shown in Figure 3. This could be used for the simulation of a range of machine faults like in the gearbox, shaft misalignments, rolling bearing damages, resonances, reciprocating mechanism effects, motor faults, pump faults, etc. In the MFS experimental setup, three-phase induction motor is mounted to the rotor that is connected to the gear box through a pulley and belt mechanism. The gear box and its assembly are illustrated in Figure 4. In the study of faults in gears, three different types of faulty pinion gears namely the chipped tooth (CT), missing tooth (MT) and worn tooth (WT) along with normal gear (or no defect i.e. ND) were used (illustrated in Figure 5). The real time data in time domain were measured using a tri-axial accelerometer (sensitivity: x-axis 100.3mV/g, y-axis 100.7mV/g, z-axis 101.4mV/g) mounted on the top of the gearbox (illustrated in Figure 6) and the data acquisition hardware. Measurements were taken for the rotational speed of 10 to 30Hz in intervals of 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000757_j.cirpj.2015.08.005-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000757_j.cirpj.2015.08.005-Figure4-1.png",
        "caption": "Fig. 4. A magnet group (M1, M2 and M3) in the master array coaxial to another magnet group (S1, S2 and S3) in the slave array in static (nominal) condition.",
        "texts": [
            " Similar arrangements are adopted both in the master and slave array. A magnet in the master module array has an opposite magnet in the slave module array, forming a pair. The magnets of each pair are nominally coaxial. The clamping contact point is at the centroid of each group where a ball transfer unit is attached (Fig. 3). A magnet attracts the opposite pole and repels the similar pole. Since these permanent magnets are axially magnetized the N-S-N-S orientation is used to get the highest possible attraction force as shown in Fig. 4. Coaxial magnets located in the master and slave array have their own magnetic attraction forces between them (Fig. 4). However, the other two magnets of the same group have an influence on that force as they are closely located. This was accounted for with the introduction of a force ratio representing the array contact force enhancement by Vokoun and Beleggia [8]. Furthermore, the member magnets of the other groups present in the array have influence on each other. However, as will be shown in the results section, such effects are negligible, due to the significant distance between the groups, and so will be ignored in the calculation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001491_1.4040809-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001491_1.4040809-Figure6-1.png",
        "caption": "Figure 6. Cross-section of the zero pre-swirl ring [7]",
        "texts": [
            " For the measurement uncertainties of the turbine flowmeter, the Coriolis flowmeter, and other instrumentations, please see Fluid out Bearing pedestal Fluid out Fluid out Fluid out Vertical stiffenerExhaust chamber support Back labyrinth seal Test seal Hydrostatic bearing ExhaustExhaust Shaker Stinger High speed coupling Coupling cover Pitch stabilizer Rotor Bearing oil inlet Bearing oil inlet Acc ep te d Man us cr ip t N ot C op ye di te d Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/ on 08/10/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use GTP-17-1637, Zhang 3 Copyright \u00a9 2018 by ASME Zhang et al. [7]. Zhang et al. [7] also describe the method to calculate the LVF at the seal inlet. During a test, the test fluid flows through a zero pre-swirl ring before entering the test seals, as shown in Fig. 5. Figure 6 shows the zero pre-swirl ring\u2019s cross-section. The zero preswirl ring guides the test fluid radially inwards, aiming to produce a flow with minimum circumferential velocity at the seal inlet. The measured inlet pre-swirl ratio u0(0) is close to zero (from -0.07 to 0.08). The uncertainty varies with inlet LVF and \u03c9 with the maximum uncertainty=0.06. Childs et al. [8] describe the method to measure u0(0) using a pitot tube. Figure 5. Section view of the stator assembly [7] All tests are performed when Pi=62"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001294_tpel.2017.2782804-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001294_tpel.2017.2782804-Figure4-1.png",
        "caption": "Fig. 4. 2-D mesh model of a 8 pole six-phase IPMSM.",
        "texts": [
            " Note further that Ld1, Lq1 and Ld2, Lq2 are the inductances in the synchronous reference frame. They are different though a single rotor is shared. The reason is that the rotor axis is seen differently by the two coil sets. It is worthwhile to note that Lq1 \u2212Ld1 = 3L\u03b4 , whereas Lq2 \u2212Ld2 = 3 2L\u03b4 . In other words, Ld1 < Ld2 and Lq1 > Lq2. Therefore, Lq1/Ld1 > Lq2/Ld2, i.e., the saliency is larger with respect to the abc-coil than to xyz-coil. C. Inductance Determination through FEM Analysis It is possible to compute the flux linkage of a coil via FEM analysis. Fig. 4 shows a 2-D mesh model of a 8 pole six-phase IPMSM in which magnets are arranged in a V-shape in each pole. The outer diameter of stator and rotor are 278 and 171 mm, respectively. The stack length is 120 mm and air gap height is 0.8 mm. For the purpose of obtaining the inductances, special cases are considered when only ia flows. FEM calculations were done for different ia\u2019s and a specific standstill rotor position, \u03b8 = \u03c0 6 . Table. I shows the computed results. Since the current to flux relation is given by \u03bb\u2032a \u03bb\u2032b \u03bb\u2032c  =  \u03bba \u03bbb \u03bbc \u2212 \u03c8pm  cos \u03b8 cos(\u03b8 \u2212 2\u03c0 3 ) cos(\u03b8 + 2\u03c0 3 )  =  Lm + Lls \u22120"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001168_icuas.2017.7991400-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001168_icuas.2017.7991400-Figure1-1.png",
        "caption": "Fig. 1. Assembled quadcopter used in this research with the on-board smartphone on the top center of it.",
        "texts": [],
        "surrounding_texts": [
            "Keywords: Quadcopter, Altitude, Attitude, Control, Smartphone, Android-based controller, Cascade PID controllers\nI. INTRODUCTION\nCurrent smartphone processors are able to perform complex calculations such as those required in the implementation of real time control strategies. There are many ongoing research related to the possibility of using smartphones to implement control strategies, such as [1], as configuration and monitoring interfaces in control systems as seen in [2], [3], and as a tool in both education and design of control strategies seen in [4], [5]. Following this trend, in the Universidad del Valle, it was developed a smartphone-based platform for monitoring, control and communication in portable laboratories, where a controller for a pendulum based in the Lego Mindstorms EV3 platform was implemented [6].\nIn recent years, the interest in aerial robotics research has increased substantially. This is because this type of robotics offers several potential new services such as search and rescue, observation, mapping, inspection, etc. On the other hand, smartphones have become essential devices for humans and easily acquirable development tools. The interaction between these two technologies allow the\nAll the authors are with the Research Group in Industrial Control of the School of Electrical and Electronic Engineering, Universidad del Valle, Cali, Colombia (e-mail: {alejandro.astudillo, pedro.munoz, fredy.alvarez, esteban.rosero}@correounivalle.edu.co).\ndevelopment of low cost quadcopters based on an everyday item such as the smartphones, facilitating the distribution of the quadcopter control software and its implementation by other researchers.\nIn the University of Pennsylvania, in [7], was developed a quadcopter using a last generation smartphone as a flight controller and an additional processing system for imagebased positioning. The state estimation algorithms, control and planning were firstly implemented in a ODROID-XU board with additional sensors, but then, in [8], this algorithms were ported to the Qualcomm processor in the phone due to the Qualcomm colaboration in that project.\nCurrent research focuses on the development of aerial robots potentiated by the use of smartphones, as seen in [9]\u2013[12]. In the last years, computing capacity and sensor technology in smartphones has decreased in price but increased in performance. Smartphones have become an inexpensive tool capable of commanding an UAV. The challenge then, is to use smartphones as quadcopter flight controllers for autonomous flights following specific missions, taking advantage of the fact that the phones today are very powerful computers that include elements of sensing, processing and signal communication.\n978-1-5090-4494-8/17/$31.00 \u00a92017 IEEE 1447",
            "In this paper, the implementation of a quadcopter with a smartphone acting as its flight controller while using exclusively the sensors and processor in the smartphone, is shown. The controller keeps the attitude of the quadcopter stabilized while making the quadcopter to hover at an altitude reference. It is presented the detailed composition of the test platform (quadcopter) used, integrating it with its dynamic model in addition to the quadcopter altitude and attitude estimation strategies using sensor fusion algorithms.\nIn the next section, a detailed description of the hardware used in the quadcopter is presented. After that, the dynamic non-linear and linearized quadcopter model is shown in Section III. In Section IV, all the aspects related to the altitude and attitude cascade controllers design are presented. Then, in Section V the sensor fusion algorithm that is used to estimate the altitude of the quadcopter, and how the orientation of the system is estimated, is described. In Section VI, the implementation in Android of all the necessary software with its results is provided. Finally, Section VII concludes this paper taking into account the results of this research and the future work to be done.\nIn this section, it is detailed the hardware used to build the quadcopter. To control the quadcopter it is necessary to sense variables such as orientation and position, and define control signals that will command the actuators. This tasks are done completely by the smartphone. The control signal goes from the smartphone, through an Arduino Mega ADK that acts as a gateway, and to the Electronic Speed Controllers (ESCs) that will set the motors speed depending on the control signal they receive.\nThe Nexus 4 is a smartphone developed by Google and assembled by LG, released in 2012 with Android 4.2 OS. This phone has a quad-core Qualcomm Snapdragon APQ8064 1.5 GHz processor, 2 GB of RAM memory and a total mass of 139 g. Its instrumentation includes accelerometers, gyroscopes, magnetometers, GNSS and\nbarometer. It handles Bluetooth 4.0, WiFi, USB and GSM connectivities.\nAll the sensorial, estimation and control algorithms are executed in the Nexus 4 Smartphone, and it sends the control signals to the Arduino Mega ADK through an USB cable.\nThe Arduino Mega ADK is a development board based on the Atmel 8-bit AVR RISC-based ATmega2560 microcontroller. The main use advantage of this board is that it handles a MAX3421E USB host chip which let us connect this board as a USB host to a Smartphone, while using the USB Host Library for Arduino. This board receives the control signals through its USB port and translate them into PWM signals that will command each of the four ESCs.\nTo power all the hardware that makes up this quadcopter, it is used a 3 cells lithium-ion polymer battery with a nominal current capacity of 4500 mAh and a nominal voltage of 11.1 V. This battery can deliver up to 20 times its nominal current, so it can deliver 90 A continuously. As seen in Fig. 2 the battery is directly connected to the Arduino Mega ADK and to the high power supply of all the four ESCs.\nTo vary the speed of a R/C-type brushless motor it is necessary to use an ESC. This controllers receive a PWM signal, a low power supply and a high power supply, and output a three-phase electric power directly to the brushless motor. Four 30 A Simonk ESCs where used. Each of the ESCs receives the PWM signal and the low power supply from the Arduino Mega ADK board, while the high power supply is received directly from the LiPo battery.\nThe brushless motors are the actuators in any multirotor. The speed of each motor is controlled by the three-phase electric signal sent by the ESCs. In this quadcopter we used four DJI 2212 920 KV motors with 10\u201d length and 4.5\u201d pitch propellers.\nIn this section, the derivation of the quadcopter dynamic model is presented. First, there are shown the equations of motion that describe precisely the way the quadcopter is affected on its angular and linear accelerations. Then, a linearized model where just the altitude and attitude dynamics are taken into account is got."
        ]
    },
    {
        "image_filename": "designv11_13_0001520_0954409718789531-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001520_0954409718789531-Figure5-1.png",
        "caption": "Figure 5. Free body diagram of the wheelset.",
        "texts": [
            " _Zbl _Ybl \" # \u00bc _Zb _Yb \" # \u00fe lp 0 0 h0 0 0 _ b _ b _ b 2 64 3 75 \u00f015\u00de _Zbr _Ybr \" # \u00bc _Zb _Yb \" # \u00fe lp 0 0 h0 0 0 _ b _ b _ b 2 64 3 75 \u00f016\u00de _Zbfl _Ybfr \" # \u00bc _Zb _Yb \" # \u00fe lp lt 0 h0 0 lt _ b _ b _ b 2 64 3 75 \u00f017\u00de _Zbfr _Ybfr \" # \u00bc _Zb _Yb \" # \u00fe lp lt 0 hg 0 lt _ b _ b _ b 2 64 3 75 \u00f018\u00de _Zbrl _Ybrl \" # \u00bc _Zb _Yb \" # \u00fe lp lt 0 hg 0 lt _ b _ b _ b 2 64 3 75 \u00f019\u00de _Zbrr _Ybrr \" # \u00bc _Zb _Yb \" # \u00fe lp lt 0 hg 0 lt _ b _ b _ b 2 64 3 75 \u00f020\u00de A bond graph fragment of the bogie sub-system is shown in Figure 3 and is denoted as LB. The model considers the vertical, lateral, roll, pitch and yaw motion of the bogie. Bond graph modelling of the wheelset sub-system The wheelset mainly provides support to the vehicle. The interconnection between the wheelsets and bogie frames are through the primary suspensions. A conical wheelset as a rigid body and a flexible rail having a knife edge (Figure 4) is considered in this work. The free body diagram of the wheelset is shown in Figure 5. The governing equations of motion for the wheelsets are given by equations (21) to (24). Mw \u20acyw \u00bc Mwg se \u00fe MwV 2 Rt \u00fe FLcry \u00fe FRcry \u00feNLy \u00feNRy Ffc \u00fe Fwsy \u00f021\u00de Mw \u20aczw \u00bc Mwg MwV 2 se Rt \u00fe FLcrz \u00fe FRcrz \u00feNLz \u00feNRz \u00fe Fwsz \u00f022\u00de Iwx \u20ac w \u00bc IwyV _ w r0 IwyV 2 Rtr0 \u00fe RRyFRcrz RRzFRcry \u00fe RLyFLcrz RLzFLcry \u00fe RLyNLz \u00fe RRyNRz RRzNRy RLzNLy \u00feMLcrx \u00feMRcrx \u00feMwsx \u00f023\u00de Iwz \u20ac w \u00bc IwyV _ w r0 \u00fe RRxFRcry RRyFRcrx \u00fe RLxFLcry RLyFLcrx \u00feMLcrz \u00feMRcrz \u00feMwsz \u00f024\u00de where Fwsy, Fwsz, Mwsx, Mwsy and Mwsz denote the suspension forces and moments on the wheelset in the respective directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003714_s12008-018-0520-6-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003714_s12008-018-0520-6-Figure6-1.png",
        "caption": "Fig. 6 Simulation with radiation and convection heat losses",
        "texts": [],
        "surrounding_texts": [
            "The developed model was first utilized to examine the effect of convection heat transfer during theAMprocess. The simulation analysiswas carried out for two cases,with andwithout convection heat losses, while all other process parameters were kept unchanged. Figures 6 and 7 illustrate the effect of convective boundary condition on the temperature variation within the part for a laser speed of \u03c0 mm/s. The predictions are obtained in each case for the time steps after activating the layers 1, 2, 3 and 4 at the time instant, when the laser beam reaches the middle of the corresponding layer. The results indicated that the absolute values of the temperature difference between the two cases increase with time. The computed temperature differences lie in the interval between 15 and 30 \u00b0C and reach a maximum relative value of approximately 3%. Subsequently, neglecting the convection losses would not lead to significant errors. Figure 8 shows the temperature variation along the height of the part for both cases. This figure clearly shows the insignificance of convective boundary condition on the results (h 30 W/m2 K). Figure 9 illustrates the thermal cycles at points A, B and C corresponding to the substrate and the two deposited layers 2 and 4, respectively. The peaks indicate the time instants at which the laser beam passes through the pre-defined locations.At themid-point of the 2nd layer, the initial temperature peak is approximately 1750 \u00b0C. After 1.5 s, the heat is conducted away and the temperature decreases to a value of about 540 \u00b0C. The mid-point of layer 4 has experienced a similar thermal cycle, where the maximum temperature has a value of 2113 \u00b0C. The temperature peaks at points B and C indicate that the melting point of the material has been reached. An interesting observation is noticed in Fig. 8 that after the deposition of the fourth layer, the mid-point temperature of the 2nd layer reached 1650 \u00b0C. This aging and tempering effects may affect the mechanical properties of the part including the residual stresses and its strength. Figure 10 demonstrates the effect of varying the speed of the laser beam on the temperature distribution of the part after completing building the last layer (4th layer). It can be seen from Fig. 10 that as the velocity of the laser beam increases, the temperature of the part decreases comparedwith the case of lower velocity. This is can be attributed to less time of exposure to the laser beam. Figure 11 shows the temperature variation along the height of the part for various laser travel velocities. The plotted data was taken at the time instant after the 4th layer was deposited. The corresponding maximum temperature were about 2105 \u00b0C, 2010 \u00b0C, 1700 \u00b0C and 1510 \u00b0C for tested scan speed of \u03c0 mm/s, 2\u03c0 mm/s, 4\u03c0 mm/s and 6\u03c0 mm/s, respectively. Figure 10 clearly shows that the temperature along the height of the part is the lowest for the higher velocity (6\u03c0 mm/s) and this is due to the shorter interaction time between the laser beam and the material layer during the scanning process. However, the lowest travel velocity exhibits themaximum temperature."
        ]
    },
    {
        "image_filename": "designv11_13_0002737_9781118899076-Figure3.11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002737_9781118899076-Figure3.11-1.png",
        "caption": "FIGURE 3.11 (a) The sealed tissue layer is placed between the two compartments, and ions moving across can be measured by radioisotopes or by means of voltage clamp or current clamp techniques [62]. (b) The isolated sweat gland impaled on microcapillary is used to measure ion transport during sweat production [77].",
        "texts": [
            " From an electrochemical point of view, while CFTR is called the chloride channel, it is in fact an anion channel, with low selectivity coefficient Cl\u2212/HCO3 \u2212 = 4 [61]. Is it possible that CFTR channel is responsible for chloride, bicarbonate, and water transport in the same time? The study of epithelial cell layer goes back to Ussing\u2019s experiments on the water transport across the frog skin [62]. The sealed layer was placed between the two compartments, and the ions moving across the skin were measured by radioisotopes. The other possibility was the use of voltage clamp or current clamp technique (Fig. 3.11a). The most common is the use of voltage clamped at zero (short-circuited) between apical and basolateral faces of the tissue system. The current measured in such configuration is called the short-circuit current. The method has obvious limitation\u2014one cannot recognize between cations flowing inward from anions flowing outward. Moreover, only the total current of anions and cations flowing through both basolateral and apical membranes can be measured. The activators and/or blockers of channels, pumps, and transporters are used to get an insight into what is being measured and 73ION CHANNELS IN HEALTH AND PATHOLOGY which transporters are involved",
            "12) and volume control processes (Fig. 3.13). The study of the sweat gland [77] gives the best insight into the mechanisms of water and ion transport in CF. The salty taste of sweat was the first discovered manifestation of CF since hypotonic sweat chloride concentrations in CF patients exceed 75ION CHANNELS IN HEALTH AND PATHOLOGY 60 mM (and the value for non-CF people is around 20 mM). The sweat gland can be isolated from the skin and put on the tip of a capillary holding reference electrode (Fig. 3.11b). Thus, not only the short-circuit current could be measured but also the ionic composition of the sweat can be determined. The data obtained suggest that isotonic sweat is produced in the secretory coil of the sweat gland by CFTRindependent mechanism and when it passes through the reabsorptive duct, the sodium chloride is removed by concerted action of CFTR and ENaC channels transporting electroneutrally NaCl. It is not certain whether the mechanism of sweat secretion can be extended on the bronchial epithelium"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003688_042044-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003688_042044-Figure1-1.png",
        "caption": "Figure 1. Schematic of electron beam freeform fabrication (EBF) system components [5].",
        "texts": [
            " Fast development of additive manufacturing is a global challenge in modern technology progress, which determines provisions for developing and using new advanced, high-productive and competing processes. This work has been focused on developing an electron multibeam additive directed energy wire deposition process and commercial high-productive robotic equipment for manufacturing large metallic components. This process involves a layer-by-layer deposition of metal by electron beam melting of wire and obtaining then a near-net-shape component (see schematics in figure 1). The advantage of this process is its high deposition rate up to 12 kg/h [1] which is unachievable with other additive processes, for example with a powder additive process [2]. Also it allows making large up to 5000 mm size fully dense and structurally homogeneous components from both refractory and heat-resistant alloys [3]. Extra feature of this process is a feasibility of simultaneous deposition of dissimilar metals and thus forming a composite structure inside a vacuum chamber [4]. Therefore, it excludes any oxidizing of the component"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002331_1350650114559997-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002331_1350650114559997-Figure8-1.png",
        "caption": "Figure 8. Test bearing inner raceway showing EDM notch.",
        "texts": [
            " A beta ratio of 1000 represents a particle removal efficiency of approximately 99.9% for the stated size.) A strain gage was fitted to determine the load applied to the test bearing. Angular contact bearings (details in Table 3) were used in this trial since the test rig provides pure axial load to the test bearing. The load was applied using a large threaded load screw that contacts the outer race of the test bearing. Each test bearing had a notch (2mm long 250mm wide 100 mm deep) electric discharge machined (EDM) into the inner raceway (Figure 8) as a stress raiser to ensure a RCF fault initiated at one location and within a reasonable test time. Additionally the load used for the test exceeded the maximum axial load rating for these bearings to further accelerate the failure; typically the load applied was in the range of 6\u20138 kN. Wear debris data was collected using the MetalSCANTM software provided with this particular inductive wear debris sensor. Once the particle count had reached the \u2018\u2018Terminate\u2019\u2019 limit, the rig was shut down and the bearing removed for at Gebze Yuksek Teknoloji Enstitu on December 20, 2014pij"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000049_robio.2013.6739433-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000049_robio.2013.6739433-Figure4-1.png",
        "caption": "Fig. 4. Illustrations of typical swimming patterns designed for the turtle-like underwater robot. (a) Forward and backward swimming; (b) ascending and submerging; (c) pitching; (d) rolling; (e) rolling motion while swimming forward; (f) rolling motion while staying in situ.",
        "texts": [
            " The maneuvers can be realized by the composition of appropriate angular offset, vibration amplitude and frequency. Specially, the force direction can be regulated through adjusting the angular offset of the flipper, while the magnitude of the force can be regulated through adjusting the vibration amplitude and vibration frequency. In this article, for the convenient of control, we set the amplitude as a fixed value, and introduce the vibration angular velocity as the integral effect of the vibration amplitude and frequency. Forward and backward swimming (Fig. 4a). The vehicle swims forward with all its foils oscillating synchronously around the horizontal plane by setting all the offsets iX to zero, while swimming backward with all the foils oscillating synchronously in the contrary direction by setting all offsets to 180\u00b0. As to swimming in a straight line, all the phase biases ij are set to 0\u00b0. Ascending and submerging motion (Fig. 4b). The vehicle swims upwards (or down) by regulating the synchronized oscillations of all foils around the vertical plane by setting all offsets X i to / 2 ( / 2 ), at the same time the phase biases ij are set to 0\u00b0. Pitching motion (Fig. 4c). The downthrust of the vehicle can be obtained by setting ( , )front backX X    , while the upthrust can be obtained by setting ( , )front backX X    . Thereinto ( 0, 0)front backX X  and 0 90    . Rolling motion (Fig. 4d). The vehicle can roll in a clockwise direction (or anticlockwise) by setting the offsets of the left pair foils leftX and the right pair foils rightX to / 2 and / 2 (or ( / 2 and / 2 )) respectively. Yawing motion while swimming forward (Fig. 4e). While swimming forward, the vehicle can yaw to the left (or right) by setting the vibration amplitude and vibration frequency of the foils on the right side larger (or smaller) than the left. Yawing motion while staying in situ (Fig. 4f). While staying still, the vehicle can yaw to the left (or the right) by setting the offsets of the left pair foils leftX and the right pair foils rightX to  and 0 (or ( 0 and )) respectively. D. Controller design At present study, the vehicle is operated by remote operating handle (F510, Logitech). As shown in Figure 5, an upper monitor (any laptop with the relative software) is used to recognize the operation signals of the handle and generate motion commands to the vehicle. By calculating the navigation information, the monitor also used to display the motion and attitude status of the vehicle in real time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001814_gt2015-43935-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001814_gt2015-43935-Figure8-1.png",
        "caption": "Figure 8. CAD image of rig used to measure pressure drop and heat transfer in additively manufactured test coupons.",
        "texts": [
            " However, as shown in Table 2, the teardrop and diamond shaped channels are smoother than the cylindrical shaped channels. This can be explained by the fusing of the small upper surfaces in the diamond and teardrop coupons. This fusing effectively smoothed some of the roughness features that would have been present on the each individual surface. A test rig was built to collect pressure drop and heat transfer measurements of each of the test coupons using a similar design to that already presented by Weaver et al [20] and Stimpson et al [18]. The rig, shown in Figure 8, was built with a smooth contraction chamber that supplied uniform velocity air to the coupon inlet. The exit expansion chamber was made identical to the inlet for simplicity of fabrication. The coupon was mounted between the inlet and exit pieces and sealed with rubber gaskets between the mating surfaces. Pressure taps were installed upstream of the inlet contraction and downstream of the exit expansion to measure the pressure drop across the coupon. This pressure drop was modified to account for expansion losses associated with a sharp exit into a large reservoir"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001016_978-3-319-30808-1_177-1-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001016_978-3-319-30808-1_177-1-Figure1-1.png",
        "caption": "Fig. 1 Model of a pendulum. Symbols: M is pendulum mass, \u03b8 is angle relative to vertical, L is distance from the suspension point to the center of mass, I is moment of inertia about the suspension point, g is gravitational acceleration, \u03c4 is a torque acting on the pendulum",
        "texts": [
            " For ease of computer implementation, most investigators make use of the fact that any higher-order differential equation may be expressed as a system of simultaneous first-order equations. To provide a concrete example, consider a physical pendulum of mass M, where \u03b8 is the angle relative to vertical, L is the distance from the suspension point to the center of mass, I is the moment of inertia about the suspension point, g is the gravitational acceleration, and \u03c4 is a torque acting on the pendulum (Fig. 1). This system has the following equation of motion \u20ac\u03b8 \u00bc MgL sin \u03b8 \u00fe \u03c4\u00f0 \u00de=I (4) and we can define \u03c4 \u00bc u\u03c40 (5) such that u is the control variable, bounded between 1 and 1, that scales \u03c4 between its maximum negative (clockwise) and positive (counterclockwise) magnitude, \u03c40. If we introduce a new variable \u03c9, where \u03c9 \u00bc _\u03b8 , then Eq. 4 may be rewritten as the following system of first-order differential equations _\u03c9 \u00bc MgL sin \u03b8 \u00fe \u03c4\u00f0 \u00de=I (6) _\u03b8 \u00bc \u03c9: (7) In this example, the state equations are given by Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002218_j.triboint.2014.04.024-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002218_j.triboint.2014.04.024-Figure15-1.png",
        "caption": "Fig. 15. Dimensionless cases.",
        "texts": [
            " In order to maintain high computation speeds, the results presented in the previous sections simulated the syy evolution from a plane strain to a plain stress state towards the free boundary with a constant radius curved transition set to \u03b3C, with \u03b3\u00bc15. The results showed that this assumption is suitable for the dimensionless cases of Table 1. To conclusively verify the acceptability of this approach, the FEM calculations presented below examine eight new dimensionless configurations taken outside the domain defined by the factor range of Table 1. Fig. 15 illustrates the study domain. Since Lr corresponds to the coincident or noncoincident end conditions, this factor remains fixed at the same values Lr\u00bc1 or 1.4. The first four cases (1\u20134) evaluate the slenderness influence with Sl\u00bc0.1 and 4, while W is maintained at a middomain load of 4.162 10 5, leading to a contact pressure of 400 MPa. The following four cases examine the load influence; W is set to 2.602 10 6 and 2.602 10 4 to generate contact pressures of 100 MPa and 1000 MPa, while Sl is fixed at the middomain value of 1.1. Fig. 15 presents the studied configuration, where points 1\u20138 correspond to the following designations: case 1: S400C, case 2: L400C, case 3: S400NC, case 4: L400NC, case 5: M100C, case 6: M1000C, case 7: M100NC, and case 8: M1000NC. The resulting syy distributions established along the contact lines for both cases of Table 1 and the eight new cases are drawn in the graphs in Fig. 16. This figure clearly demonstrates that, following an arc-form reduction, the syy value reaches more than 90% of its central amplitude on average, at a distance close to 15 times the semi-width of contact (C)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002112_j.rcim.2014.09.003-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002112_j.rcim.2014.09.003-Figure6-1.png",
        "caption": "Fig. 6. Tensile tests for a round of beef (a) a wooden support for sample\u2019s ends freezing, (b) freezing in liquid nitrogen, (c) failure of the test piece, (d) stress evolution based deformations.",
        "texts": [
            " From this finding, it is necessary to conduct a series of tests on several samples of muscles with the aim of ensuring a better representation of the mechanical behavior of muscles. The objective is to characterize the handled muscles based on compression tests (until rupture of the sample) and tensile tests (up to 80% deformation). This allows identifying different rheological properties involved in muscles mechanical models. Tests have defined the evolution of constraints according to the deformation (Fig. 6). Consequently, for each range of meat deformation, the elasticity modulus is defined. Tests were made on three directions: the longitudinal direction, which is the direction of meat fibers, transversal tests and 451 tests. For each test, five cubic samples of 10 mm side are used and a low strain rate (in the order of 0:05% s 1) is considered to minimize the viscous effects. For each test, stress curves are averaged before parameter identification. These tests are elaborated in collaboration with INRA1. We notice that for tensile tests, it is difficult to use the conventional clamping techniques to attach the ends of the samples. An alternative to this problem is to use a cryogenic cell [27] whose principle is to freeze the ends of the sample to be studied while the temperature at its center remains unchanged (Fig. 6). The linear tensor-mass model was introduced by Cotin et al. [12] and extended by Picinbono et al. [13] to non-linear elasticity based on a St Venant Kirchoff model (hyperelastic materials). For both linear and non-linear formulations, the elastic energy for an isotropic material is given by: W \u00bc \u03bb 2 tr E\u00f0 \u00de2\u00fe m tr E2 \u00f01\u00de where \u03bb and m are the Lam\u00e9 coefficients characterizing the material stiffness. In the linear case, tr\u00f0E\u00de and tr\u00f0E2 \u00de represent the principal invariants of the linearized Green\u2014St"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001020_j.comgeo.2017.03.001-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001020_j.comgeo.2017.03.001-Figure3-1.png",
        "caption": "Fig. 3. The common development shown in [6]. (a) It folds into a box of size 1 \u00d7 2 \u00d7 4 and (b) it also folds into a box of size \u221a 2 \u00d7 \u221a 2 \u00d7 3 \u221a 2.",
        "texts": [
            " We remark that we cannot use the same way as one for area 22 shown in [1] since it takes too huge memory even on a supercomputer. Therefore, we use a hybrid search of the breadth first search and the depth first search. Our first result is the number of common developments of two boxes of size 1 \u00d7 1 \u00d7 7 and 1 \u00d7 3 \u00d7 3, which is 1,080. Based on the obtained common developments, we next change our scheme. In [6], they also considered folding along 45 degree lines, and showed that there was a polygon that folded into two boxes of size 1 \u00d7 2 \u00d7 4 and \u221a 2 \u00d7\u221a 2 \u00d7 3 \u221a 2 (Fig. 3). In this context, we can observe that the area 30 may admit to fold into another box of size \u221a 5 \u00d7 \u221a 5 \u00d7 \u221a 5 by folding along the diagonal lines of rectangles of size 1 \u00d7 2. This idea leads us to the problem that asks if there exist common developments of three boxes of size 1 \u00d7 1 \u00d7 7, 1 \u00d7 3 \u00d7 3, and \u221a 5 \u00d7 \u221a 5 \u00d7 \u221a 5 among these 1,080 common developments of two boxes of size 1 \u00d7 1 \u00d7 7 and 1 \u00d7 3 \u00d7 3. We remark that this is a special case of the development/folding problem mentioned above. In our case, P is one of the 1,080 polygons that consist of 30 unit squares, and Q is the cube of size \u221a 5 \u00d7 \u221a 5 \u00d7 \u221a 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001146_0954406217722380-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001146_0954406217722380-Figure1-1.png",
        "caption": "Figure 1. Photo of scaled test rig: (a) test rig; (b) setting unbalance.",
        "texts": [
            " In order to know well the cage motion, the real boundary conditions of the bearing should be considered. A scaled test rig for one original rotor system in aero-engines is designed and set up, satisfying the geometric and dynamic similarities proposed in Wang et al.33 The scaled test rig consists of a motor, a disc, and a shaft, which is supported by two support assemblies and driven by the motor through a flexible coupling. And the tested bearing is an angular contact ball bearing of 7013AC located on support B1, as shown in Figure 1(a). The disc has 12 equally spaced threaded holes, as shown in Figure 1(b), helpful in adjusting the rotor unbalance mass. The geometrical parameters of the rotor system and bearings are listed in Tables 1 and 2. As is well known, unbalance is always present in all rotating machineries in reality. So it is necessary to determine the initial state firstly. The equivalent rotor assembly is dynamically balanced to ISO G6.3 standard, which can be defined as the initial state. Besides, the appropriate axial load applied to the tested bearing should be constant. The outer ring of the tested bearing is cut specially to form four grooves at 90 apart, which are of 20mm width and 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002432_978-3-319-14705-5_5-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002432_978-3-319-14705-5_5-Figure1-1.png",
        "caption": "Fig. 1 Manipulator motion in the presence of some obstacles",
        "texts": [
            " Then the secondary task is defined by some motion xt = ft (q) like in the case of obstacle avoidance, the velocity \u03d5\u0307 can be defined as \u03d5\u0307 = J+ x\u0307 (7) Another possibility is to define \u03d5\u0307 as \u03d5\u0307 = Kp\u2207 p, (8) where, p is a function representing the desired performance criterion, \u2207 p is the gradient of p, and Kp is a gain. So, using (8) the optimization of p can be achieved. The obstacle-avoidance problem usually defines how to control the manipulator in order to track the desired end-effector trajectory while simultaneously ensuring that no part of the manipulator collides with any obstacle in the workspace of the manipulator. To avoid any possible obstacles the manipulator has to move away from them into a configuration where the distance between them becomes larger, as shown in Fig. 1. Reconfiguration of themanipulator without changing themotion of the endeffector is only possible if the manipulator has redundant DOFs. Note that in some cases it is possible that the redundant manipulator cannot avoid an obstacle, because it might be in a configuration where the avoiding motion in the desired direction is not feasible. Having a high degree of redundancy reduces the chance of getting into a such configuration, especially if the manipulator is working in an environment that has many potential collisions with obstacles. Usually, the basic strategy for obstacle avoidance is to identify the points on the robotic arm that are near obstacles and then assign to them the motion component that moves those points away from the obstacle, as shown in Fig. 1. The robot motion (configuration) is changed if at least one part of the robot is at a critical distance from an obstacle. We denote the obstacles that are closer to the critical distance as the active obstacles and the corresponding closest points on the body of the manipulator as the critical points. For industrial robots it is usually assumed that the motion of the end-effector is not disturbed by any obstacle. If such a situation occurs, either the task execution has to be interrupted and the higher-level path planning has to recalculate the desired motion of the end-effector or if the path-tracking accuracy is not important the control algorithms that move the end-effector around obstacles on-line can be used",
            " There is a variety of sensor systems that can be used for such obstacle detection. In many cases a vision system is used to detect obstacles. Another possibility is offered by tactile sensors, like artificial skin, which can detect the obstacle only if they touch it, or by proximity sensors, which can sense the presence of an obstacle in the neighborhood. The basic strategy for obstacle avoidance considers the obstacle-avoidance problem at the kinematic level. We denote x\u0307e as the desired velocity of the end-effector, and Ao as the critical point on the obstacle (see Fig. 1). To avoid a possible collision, one possibility is to assign a velocity to Ao such that it would move the manipulator away from the obstacle, as proposed in [17]. Here, the motion of the end-effector and the critical point can be defined as Jq\u0307 = x\u0307e Jo q\u0307 = x\u0307o (9) where Jo is a Jacobianmatrix associated with the point Ao. In the following, different possibilities for finding the solution for both equations will be presented. Let x\u0307 in (5) be equal to x\u0307e. Then, by combining (5) and (9) we obtain \u03d5\u0307 = (JoN)#(x\u0307o \u2212 JoJ# x\u0307e) (10) Using \u03d5\u0307 in (5) gives the final solution for q\u0307 in the form q\u0307 = J# x\u0307 + (JoN)#(x\u0307o \u2212 JoJ# x\u0307e) (11) Note that N is both hermitian and idempotent [4, 17]",
            " Hence, the rank of JoN is one, and the pseudo-inverse (JoN)# does not give a feasible solution, at least the desired avoiding velocity x\u0307o cannot be achieved. On the other hand, as the obstacle-avoidance strategy only requires motion in the direction of the line connecting the critical pointwith the closest point on the obstacle, this is a one-dimensional constraint for which only one degree of redundancy is needed. Therefore,we propose using a reduced operational space [39] for the obstacle avoidance and define the Jacobian Jo as follows. Let do be the vector connecting the closest points on the obstacle and the manipulator (see Fig. 1) and let the operational space in Ao be defined as one-dimensional space in the direction of do. Then, the Jacobian that relates the joint-space velocities q\u0307 and the velocity in the direction of do can be calculated as Jdo = nT o Jo (14) where Jo is the Jacobian defined in the Cartesian space and no is the unit vector in the direction of do, no = do\u2016do\u2016 . Now, the dimension of the matrix Jdo is 1 \u00d7 n, and the velocities x\u0307o and Jdo J# x\u0307e become scalars. Consequently, the computation of (Jdo N)# is also much faster [33, 35, 39]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001654_012003-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001654_012003-Figure6-1.png",
        "caption": "Figure 6. Roughness measurement line.",
        "texts": [
            "1088/1757-899X/441/1/012003 As can be seen from hardness analysis, the specimens hardness reaches 35-38 HRC units. The heat treatment conditions: heating, holding at a temperature of 860\u00b0C for an hour, oil quenching, tempering of 180\u00b0C for an hour. Based on the measurement results, the specimens hardness after heat treatment reaches 57-61 HRC values at the following SLM process parameters: laser power 125 W, scan speed 550-700 mm/s and scanning step 100 microns. A study of the roughness of the samples was performed on the samples shown in figure 6. The results of measuring the SLM samples roughness are shown in figure 7. ISPCIET\u20192018 IOP Conf. Series: Materials Science and Engineering441 (2018) 012003 IOP Publishing doi:10.1088/1757-899X/441/1/012003 Based on the conducted researches, it can be concluded that the smallest roughness measured along the scanning direction (figure 7a) is achieved at scan speeds in the range of 400 and 700 mm/s and a laser power range from 125 to 180 W. The lowest roughness is observed at a scanning step of 100 \u03bcm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000513_smc.2014.6974513-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000513_smc.2014.6974513-Figure3-1.png",
        "caption": "Figure 3 The definition of the arm angle",
        "texts": [
            " In addition, the origin of the 3th frame which is located on the elbow (joint 4) is denoted by the point E. As shown in Figure 2, the plane SEW is defined as the arm plane [8]. Given a reference plane, the arm angle \u03c8 , i.e. the angle between the reference plane and the arm plane, is defined to represent the self-motion [8]. Without loss of generality, the line SW is used as a line of the reference plane. Then only one more point Q is needed to determine the reference plane, which is then represented as SQW. The relationship of the reference plane, the arm plane and the arm angle is shown in Figure 3. For the convenience of discussion, we define some vectors as follows: w : A vector from the point S to the point W; e : A vector from the point S to the point E; V : The vector from the point S to the point Q in the reference plane; k p, : The vectors which are perpendicular to w in the B. Resolutions of Joint Angles 1) Determining the Elbow Joint Angle In order to separate the elbow joint angle, we project the arm plane to the plane which is perpendicular to the axis of joint 4. The projected plane is denoted as S E W\u22a5 \u22a5 (see Figure 4)",
            " If \u03c8 is given, the orientation of the 3th frame relative to the 0th frame can be given by 0 0 0 0 3 3 \u03c8 \u03c8 ==R R R (7) From the definition of the rotation matrix, we can have ( )0 0 1 2 3 1 2 3 1 2 3, ,f \u03b8 \u03b8 \u03b8= =R R R R (8) Resolving the equations (7) and (8) , we get two sets of solutions about 1 2 3, ,\u03b8 \u03b8 \u03b8 . 3) Determining the Wrist Joint Angles The attitude 4th frame with respect to the 7th frame is given by: ( )( )7 0 0 0 0 3 4 7 3 4 \u03c8 \u03c8 == TR R R R R (9) On the other hand, ( ) ( )T7 4 5 6 4 5 6 7 5 6 7, ,g \u03b8 \u03b8 \u03b8= =R R R R (10) According to (9) and (10), we can get two groups 5 6 7, ,\u03b8 \u03b8 \u03b8 C. Existing Problems for Previous Methods 1) Algorithm Singularity By the definition of the reference plane and the arm angle in Figure 3, we notice that when the point W is located in the straight line which includes the vector V and passes through the point S, the reference plane degenerates to a straight line. This situation is called algorithm singularity. The literature [8] described this problem, but did not supplied feasible methods to avoid it. 2) Only Four Solutions are Determined for A Given Armangle The resolution process of the original arm angle parameterization method can only solve 4 sets of solutions. In fact, when the redundancy parameter is selected, a S-R-S manipulator should have 8 sets of solutions at least",
            " The angles between the arm plane SEW and two reference planes are defined as two arm-angles. The wrist point W cannot lie on the two different lines at the same time, showing that there is always an effective arm angle at least. It can be used to parameterize the self-motion. Without loss of generality, to simplify the calculation, two orthogonal vector -- the z0-axis and x0-axis are chosen as the reference vectors. The corresponding arm angles are denoted by z\u03c8 and x\u03c8 respectively, and called dual arm-angles. A. Dual Arm-angle z\u03c8 and x\u03c8 According to Figure 3 and the previous definitions, the projection of the vector e on the vector w is given by \u02c6 \u02c6 \u02c6, T wd = w(w e) w = w (12) The unit vector in the arm plane which is vertical to the vector w is given by ( )\u02c6 \u02c6 \u02c6,= = \u2212 = \u2212 Tpp p e d I ww e p (13) The unit vector in the reference plane which is vertical to the vector w is given by ( ) ( )\u02c6 \u00d7 \u00d7= = \u00d7 \u00d7 = \u00d7 \u00d7 =, w V w w w V w w Vnnn n (14) Considering (13) and (14), and the properties of dot product and cross product, we can have ( ) ( ) \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 c s \u03c8 \u03c8 \u00d7 = = \u00d7 = T T T n p w n p w n p (15) Therefore, ( ) ( )( )\u02c6atan2 , atan2 ,s c\u03c8 \u03c8\u03c8 \u00d7= = T Tw V p V p (16) To get dual arm-angle, we establish the same attitude axes with the base frame at the point S shown in Figure 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003777_s40799-019-00308-0-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003777_s40799-019-00308-0-Figure7-1.png",
        "caption": "Fig. 7 Sensors installation: (a) accelerometers and temperature sensors installation; (b) torque sensors installation",
        "texts": [
            " For the difference of service cycle between the two experiments, we deem that working conditions such as ambient temperature and loading process may affect its health condition. The ambient temperature of Experiment B is relatively higher than that of Experiment A. It could speed up the melting of grease, thus bad lubrication conditions inevitably lead to aggravation of damage. Three categories sensors: accelerometers, temperature and torque sensors are applied in this study. For vibration signal, four accelerometers installed at the fixed ring of slewing bearings are shown in Fig. 7(a) below and each of them is arranged every 90 degree in order to monitor the degradation process contained in vibration information more comprehensively. Temperature sensors (PT100, a platinum thermistor having a resistance value of 100 \u03a9 at the temperature of 0 \u00b0C) are also installed at the oil fill hole presented in Fig. 7(a). Torque measurement relies on the torque measuring instrument which is fixed on the drive assembly by couplings. Experimental slewing bearings are driven by the small gearing mounted on the drive assembly and powered by the hydraulic motor. Details can be found in Fig. 7(b). Experiment A & B lasted about 250 and 150 h respectively and experimental slewing bearings got stuck finally. Next, we checked the damaged parts of slewing bearings shown in Fig. 8 below and intuitively discovered the burnt cages and Table 1 Three domain candidate features Feature Equations Feature Equations Feature Equations Maximum Xmax =max {| xn| } Kurtosis X k \u00bc 1 n \u2211 n i\u00bc1 xn\u2212x \u03c3 4 Center of gravity frequency Fc \u00bc \u2211 n i\u00bc1 f ipi \u2211 n i\u00bc1 pi Variance X v \u00bc 1 n xn\u2212xn\u00f0 \u00de2 Shape factor X sf \u00bc X rms X am Mean square frequency Msf \u00bc \u2211 n i\u00bc1 f 2i pi \u2211 n i\u00bc1 pi Root mean square X rms \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffi 1 n \u2211 n i\u00bc1 x2n r Crest factor X crf \u00bc Xmax X rms RMS frequency RMSF \u00bc ffiffiffiffiffiffiffiffiffiffi MSF p Absolute mean X am \u00bc 1 n \u2211 n i\u00bc1 jxnj Impulse factor X if \u00bc Xmax Xam Frequency standard deviation VF \u00bc \u2211 n i\u00bc1 f i\u2212Fc\u00f0 \u00de2pi \u2211 n i\u00bc1 pi Root amplitude X ra \u00bc 1 n \u2211 n n\u00bc1 ffiffiffiffiffiffiffijxnj p 2 Clearance factor X clf \u00bc Xmax max X ra IMF energy X imf \u00bc \u2211 n i\u00bc1 Ci t\u00f0 \u00dej j2 rolling bodies with severe deformation with our own naked eyes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001786_9783527344758-Figure8.2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001786_9783527344758-Figure8.2-1.png",
        "caption": "Figure 8.2 Schematic illustration of a three-dimensional (a), an axisymmetric (b), and a two-dimensional (c) capsule enclosed by a membrane developing in-plane tension \ud835\udf0f and the transverse shear tension q and elastic bending moments m.",
        "texts": [
            " During the deformation of RBC and hence displacement of the membrane, the velocity is continuous across the membrane, i.e. it satisfies the no-slip condition. However, there is a jump in the interfacial tension \u0394F\u20d7 across the membrane as follows: \u0394F\u20d7 = \u0394Fnn\u20d7 + \u0394Ftt\u20d7 = \u2212dT\u20d7 dl = \u2212 d dl (\ud835\udf0f t\u20d7 + qn\u20d7) (8.1) 194 Computational Approaches in Biomedical Nano-Engineering where T\u20d7 is the membrane tension that can be decomposed into two components: the in-plane tension \ud835\udf0f and the transverse shear tension q. Here, n\u20d7 and t\u20d7 are unit vectors in the directions normal and tangential to the membrane surface, respectively (see Figure 8.2). The salient feature of the Pozrikidis formulation is that it can be used with any constitutive laws that describe the in-plane tension \ud835\udf0f and transverse shear tension q such as the Hooke (H) law, the neo-Hookean (NH) law, the Yeoh law (YE), the Skalak et al. (SK) law, and the Evans (EV) law. A detailed description of these constitutive laws is given in Dimitrakopoulos (2012). The NH constitutive law is the most widely used one due to its simplicity (Bagchi et al. 2005; Bagchi 2007; Zhang et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001321_chilecon.2017.8229691-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001321_chilecon.2017.8229691-Figure9-1.png",
        "caption": "Fig. 9. Aircraft design CAD",
        "texts": [
            "1854  Figure 8 shows the block diagram of the lift force analysis considering previous equations with pitch angle as an input and lift force as an output. IV. RESULTS Calculated values of lift, drag and pitching moment for the wing are tested on FoilSim III under following conditions: Altitude respect to sea level: 2975 m, attack angle: 7\u00b0 (approximated relative angle between multi rotor rotation to go forward and wing incidence angle), % Camber: 5.2 (Geometry of wing), chord: 0.2 m, span: 1.2 m, speed: 12.5 m/s. Lift force is 23N and drag force is 3.17N. Figure 9 shows the original design realized in a modeling software, which allows to create a detail concept of the aircraft, it shows the whole aircraft with fuselage, empennage, wing, motors and propellers. Figure 10 shows the built prototype. Figure 11 shows the flight time according to the obtained telemetry data. Li-Ion batteries have a high voltage drop at the start time that stabilizes later. To obtain the flight time, we consider the regions where the graph decreases. Based on APM Planner Log Graph, battery was active for 1975 seconds (33 minutes approximate) and flying for 1081 seconds (18 minutes)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003677_1.4042041-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003677_1.4042041-Figure2-1.png",
        "caption": "Fig. 2 Photograph of the compressor and the bearing with O-ring support",
        "texts": [],
        "surrounding_texts": [
            "Figures 1 and 2 show the structure and photograph of saturated water journal bearings implemented on a turbo compressor. The schematic of the rotor system is shown in Fig. 3. The rotor was driven by an induction motor, and its ambient pressure was lowered to the saturation point of water. Even though such auxiliary devices as a water-supplying system are not shown, they are basically the same as the previously shown apparatus [1], except that this is an operating compressor instead of a bearing test rig. The compressor was operated on the specifications listed in Table 1. This bearing was comprised of both cylindrical and tapered parts. The influence of the tapered part was ignored, because it generally produces weak thrust force and only slightly contributes to the rotor\u2019s radial dynamics. The bearings are elastically supported with deformable material (O-rings). Pressurized water is supplied to the rotor\u2019s end spaces and to the SFD: the clearance between the bearing outer cylinder and the casing\u2019s inner cylinder. The water in the SFD is discharged through a hole to the ambient space. The hole\u2019s diameter is tiny compared to the inlet hole toward the SFD so that the pressure in the SFD is kept high enough to prevent the cavitation of the water. The rotor displacements are measured with eddy current sensors at both sides of the shaft along the three directions shown in Fig. 3. The coordinate system is selected so that the Z-axis aligns to the shaft center, and the Y-axis is vertical. There is a narrow notch at 0 deg. on the surface where displacement sensors LSZ and HSZ measure. This produces a single pulse per rotation and functions as a keyphasor."
        ]
    },
    {
        "image_filename": "designv11_13_0000511_edpe.2015.7325343-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000511_edpe.2015.7325343-Figure1-1.png",
        "caption": "Fig. 1. Five-phase inverter feeding five-phase IM.",
        "texts": [
            " These higher spatial waves can increase additional losses in the machine and give rise to undesired torque pulsations. It is possible to substantially reduce the negative effects of the higher spatial harmonics and, therefore, improve the properties of the drives with multi-phase machines by suitable choice of the connection between the machine and the inverter. II. INVESTIGATED SYSTEM The investigated system consists of a five-phase induction machine fed from a five-phase voltage-source inverter. A simplified scheme of the system is shown in Fig. 1. The DClink voltage was considered constant in this study. It is possible to connect the five-phase machine to the inverter in three different ways. They are denoted as star, pentagon, and pentacle connections, Fig. 2. The output voltage of the inverter was generated by choosing appropriate combinations of switched elements. Every combination corresponds to one voltage vector of particular position and magnitude in the reference planes of the first and of the third symmetrical components of the instantaneous quantities and to a real value of the fifth symmetrical component of the instantaneous quantities [12], Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003591_j.snb.2018.09.033-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003591_j.snb.2018.09.033-Figure3-1.png",
        "caption": "Fig. 3. A) scheme of an individual glucose biofuel cell and its components. B) Assembly of 5 biofuel cells in a stack and C) photo of this stack with individual supply with glucose / O2 solution for each biofuel cell.",
        "texts": [
            " The open side of the pellet was then covered with a cellulose sheet which served as electrolyte reservoir. A PTFE foil of 127 \u03bcm thickness with a 13mm hole was placed in between the bioanode and the biocathode as a separator and to prevent any leakage. The formed bioanode and biocathode Plexiglas slices were then assembled in a stack containing 5 glucose biofuel cells, fixed with screws, and connected in series. Each glucose biofuel cell was supplied individually with glucose / O2 electrolyte solution to avoid ionic short circuit between the different bioanodes and biocathodes (Fig. 3C). Since GOx produces as side reaction H2O2 in presence of oxygen which can affect the enzymatic activity of the biocathode, the flow was directed from the biocathode to the bioanode also to reduce the oxygen concentration at the bioanode as much as possible. The used ovulation tests (Clearblue\u2122, P/N 505121/H) were bought in a pharmacy. Fig. 1A show a photo of the upside of the test holder with an eject button to remove the test stick after measurements and the display screen. To get more insight in the architecture, the ovulation test was opened revealing its internal configuration as shown in Fig",
            " With the aim to demonstrate the capability of our biofuel cell stack to supply sufficient power throughout the test, we connected 5 GOx-NQ / BOD biofuels in series, each with an OCV of 0.7 V, to provide the necessary voltage output. It should be noted that an individual single biofuel cell provides a maximum power of 1.03 \u00b1 0.05mW at 0.34 V at 5mmol L\u22121 glucose. This performance corresponds to power densities of 0.77mW cm\u22122, 3.8 mW mL\u22121, and 8.92mW g\u22121 for the bioelectrodes of 1.3 cm diameter and 2mm thickness. Each biofuel cell in the stack is individually supplied with 5mmol L\u22121 air saturated glucose solution in Mc Ilvaine buffer at pH7. Fig. 3A shows a scheme of an individual glucose biofuel cell with naphthoquinone mediated GOx as anodic biocatalyst and BOD, oriented and connected by protoporphyrin IX at the biocathode, the assembly of 5 biofuel cells, connected in series, in the stack (Fig. 3B) and a photo of the stack (Fig. 3C). This stack was connected in parallel to the ovulation test via soldering two wires to the battery connectors before removing the battery. Fig. 4 shows a photo of the experimental setup with the stack, the ovulation test without battery and a multimeter for monitoring the current during the test. We ran the experiment under identic conditions as described above by further connecting a potentiostat (not shown) for recording the voltage evolution. Before starting the test, our stack showed an OCV of 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002583_icma.2014.6885682-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002583_icma.2014.6885682-Figure2-1.png",
        "caption": "Fig. 2. Definition of variables for the yaw moment compensation algorithm",
        "texts": [
            "00 \u00a92014 IEEE Let us consider a world coordinate frame in which the z axis points up and a humanoid robot for which we have already defined some desired motion required by a main task. Also, let us assume that this motion has been previously stabilized by moving its waist horizontally so that the Zero Moment Point (ZMP) remain inside of the polygon of support. Then, the trajectory of this point may be described by rzmp/0 (t). This can be done as described in [8]. Having done this, let us denote the net moment of the ground reaction force fzmp about the ZMP, rzmp/0, by \u03c4zmp (see Fig. 2). The moment of fzmp about the origin of the world coordinate frame, \u03c40, can be calculated as \u03c40 = rzmp/0 \u00d7 fzmp + \u03c4zmp (1) In addition, we know that for the robot the following relationships between these quantities and the total linear and angular momentum about the origin of the world coordinate frame, P0 and L0, are given by P\u03070 = m\u0303g + fzmp (2) L\u03070 = rcom/0 \u00d7 m\u0303g + \u03c40 (3) Where g = [ 0 0 \u2212g ]T is the gravity vector and g is the acceleration due to gravity, whereas m\u0303 stands for the total mass of the robot and rcom/0, for the position of the center of mass (CoM) of the robot with respect to the origin of the world coordinate frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000036_j.proeng.2012.04.027-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000036_j.proeng.2012.04.027-Figure4-1.png",
        "caption": "Fig. 4. Geometric relationship between lateral force",
        "texts": [
            " The dark line shown in the figure means lateral force change and the light color line indicate the magnitude of the wake velocity fluctuation. Both fluctuations were found to fluctuate large by the order of 10 seconds. When a large fluctuation was caught in the wake velocity, an average of 3 N of lateral force fluctuation was obtained in the lateral force. The magnitude of big change of the lateral force fluctuation is possible to make the ball sway. This phenomenon indicates the lateral force occurs opposite to a vortex tail shown in Fig. 4. Change of drag as Passmore [6] investigated does not cause the erratic flight path with low frequency. Using wavelet analysis can find out a relationship of unique specification between the fluctuation of wake velocity and that of lateral force. The analysis result of lateral force fluctuation is shown in Fig. 5(a). The upper part of the figure shows captured data form and the lower part indicates the analysis result. The horizontal axis shows elapsed time and the vertical axis indicates a cycle of the lateral force component"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001956_citcon.2015.7122604-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001956_citcon.2015.7122604-Figure5-1.png",
        "caption": "Fig. 5: Axial current 1low through bearings",
        "texts": [
            " If the impedance of the return cable is too high, and if the stator grounding is poor, the current can take the path from the stator, through the bearings and through the shaft to ground in the driven machinery. Fig. 4. There is a common mode disturbance, causing current asymmetry between the three phases in the stator windings. The current sum over the stator circumference is not zero - HF flux variation is surrounding the shaft, creating a HF shaft voltage. Therefore, there is a risk for an axially flowing current through the rotor which runs through one bearing and back through the other. Fig. 5. In rolling element bearings, the rolling element is separated from the rings by a lubricant film. The lubricant film acts as a dielectric charged from the rotor. For high frequency currents, it forms a capacitor. The capacitance depends on the type of lubricant - thickener and oil type - lubricant film thickness, temperature, and viscosity of the lubricant. When the 978-1-4799-5580-0/15/$31.00 \u00a9 IEEE 2014-CIC-1045 voltage reaches the voltage breakdown limit of the lubricant, the capacitor will be discharged and a high frequency current occurs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000493_j.triboint.2014.10.005-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000493_j.triboint.2014.10.005-Figure2-1.png",
        "caption": "Fig. 2. (a) Microtextured plate (full view) with cylindrical dimple microtextures within the region from Ri \u00bc 14 mm to Ro \u00bc 19:5 mm. Total plate radius R\u00bc 20mm. (b) Microtextures imaged with 3D optical microscopy (focus variation method) have width W , periodic length L, and radial spacing LR . (c) Seven dimensional variables define the model microtexture with a Newtonian fluid. H is gap height, L is microtexture periodic length, W is microtexture width, D is microtexture depth, U is top plate velocity, \u03c1 is fluid density and \u03b7 is fluid viscosity.",
        "texts": [
            " At present, the glass bottom plate allows visual access and high alignment to the instrument axis of rotation but temperature is Contents lists available at ScienceDirect journal homepage: www.elsevier.com/locate/triboint Tribology International http://dx.doi.org/10.1016/j.triboint.2014.10.005 0301-679X/& 2014 Elsevier Ltd. All rights reserved. n Corresponding author. Tel.: \u00fe1 217 333 6532. E-mail address: ewoldt@illinois.edu (R.H. Ewoldt). not controlled. Custom-fabricated disks with microtextures (Fig. 2) are attached to the rotating component. To achieve gap control down to 20 \u03bcm, we have minimized planar misalignments and carefully calibrated the gap thickness, H. Flatness of plates is also considered and controlled. The instrument measures torque and normal force at the top geometry. The work here will focus on the torque measurements. From a gap-control perspective, this work aims to identify the correct dimensionless parameters, develop the experimental capability of a triborheometer with precise alignment and visualization capabilities, and compare computational results and gap-controlled experiments for friction reduction due to fully-wetted microtextured surfaces. (For the present study, friction refers to shear stress independent of normal force since the accurate measurement of normal force is non-trivial due to parallelism effects. As such, normal force measurement is outside the scope of the present work.) In this study, we consider circular microtextures in a deterministic layout with Newtonian fluids in hydrodynamic lubrication (Fig. 2). Seven independent parameters define the design space. With dimensionless groups, this number is reduced to four. This section defines the dimensionless parameter space, validates the choice with numerical simulations, and uses the dimensionless space to summarize related prior work. The parameterized geometry of a single texture is shown in Fig. 2c. For experiments, this is a cross-section view of a cylindrical texture. For computations, Fig. 2c defines the two-dimensional geometric model. We assume steady-state conditions with incompressible Newtonian fluid, and neglect gravity. Numerically, we solve the two-dimensional mass conservation (continuity) and momentum conservation (Navier\u2013Stokes) equations \u2207 u \u00bc 0 \u00f01\u00de \u03c1 u \u2207 u \u00bc \u2207p\u00fe\u03b7\u22072u \u00f02\u00de where u is velocity, \u03c1 is fluid density, \u03b7 is Newtonian fluid viscosity and p is pressure. We assume no-slip boundary conditions on the walls with periodic boundary conditions on the inlet and outlet. Cavitation is not considered",
            " We use the second-order upwind scheme for momentum discretization and the coupled scheme for the pressure discretization with a convergence criteria for all residuals of 10 7. The fluid is assumed to be isoviscous and constant density. The energy equation is not considered. FLUENTs, a commercial computational fluid dynamics (CFD) software, is used for all computational results. The two dependent quantities of interest are average normal pressure on the top plate, P, and average shear stress on the top plate, T. For a simple 2D periodic case these measured quantities are a function of the seven dimensional parameters shown in Fig. 2c, Measured quantities\u00bc \u00f0P; T\u00de \u00bc f \u00f0H; L;D;W ;\u03c1;\u03b7;U\u00de \u00f03\u00de where H is gap height, L is microtexture periodic length, W is microtexture width, D is microtexture depth, U is top plate velocity, \u03c1 is fluid density and \u03b7 is fluid viscosity. The Buckingham Pi Theorem reduces the complexity to four independent dimensionless variables, shown in Table 1. Two geometric ratios Dn D=W and Wn W=L define the microtexture while Hn D=H and ReH \u03c1 \u03a9RH=\u03b7 are defined by the film thickness, fluid properties, and dynamic state (velocity)",
            " Viewing the microtexture geometry design from this dimensionless perspective allows a simplified framework for future design of components. Experiments were performed on a single-head rotational rheometer (Discovery Series Hybrid Rheometer (DHR), model HR-3, TA Instruments). The instrument has a manufacturer stated torque range of 5 nN m to 200 mN m and a maximum velocity of 300 rad/s. The gap height can be controlled with a resolution of 0.1 \u03bcm. A custom-built bottom plate (Fig. 1) and custom-fabricated top plate of radius R\u00bc 20 mm (Fig. 2) were used for all tests. The fluid used is a highly refined mineral oil (Cannon Instrument Company, S60 Viscosity Standard) with a viscosity \u03b7\u00bc123 mPa s at a temperature of \u03b8\u00bc22 1C. The thermal sensitivity at 22 1C is \u2202\u03b7=\u2202\u03b8\u00bc 6:1 mPa s=K. The rheometer measures torque M, angular velocity \u03a9 and apparent gap Ha. Apparent shear viscosity \u03b7a for a parallel disk rheometer is calculated as \u03b7a \u00bc 2HaM \u03c0R4\u03a9 : \u00f07\u00de which assumes laminar simple shear flow between rotationally symmetric parallel flat disks [38]",
            " The dial indicator contacts at a radius R\u00bc20 mm and is rotated to measure the maximum peak-to-peak deviation around the plate with respect to a plane normal to the axis of rotation. The high thread count screws are adjusted until the maximum peak-topeak deviation is less than 1 \u03bcm. The microtextured plates are circular disks made of hardened stainless steel. A flat plate was also made as a control. Textures were fabricated using electric discharge machining (EDM). Microtextures are located on one side in a radial band extending from an inner radius Ri \u00bc 14 mm to an outer radius Ro \u00bc 19:5 mm (Fig. 2a). The depth D, width W, and periodicity L define the model cylindrical textures. The as-fabricated dimensions are given in Table 2, including the radial spacing, measured waviness, and measured roughness. For geometry D50.W200, the radial spacing between each ring of microtextures LR\u00bc501 \u03bcm, which is slightly larger than its circumferential periodicity L\u00bc401 \u03bcm. For geometries D300. W300, D70.W300, and D7.W300 the radial spacing LR L. The root mean square (RMS) roughness Rq and RMS waviness Lq were measured for each microtextured plate, in representative regions between microtextures",
            " Both of these effects contribute to gap offset error, so that the true gap H is larger than the apparent gap Ha from the \u201cfirst point of contact\u201d calibration. The angular misalignment may also contribute to measurement of normal forces; this can be used to identify the relative importance of misalignment to the gap offset error [44,45]. In our experiments, angular misalignment is negligible compared to the parallel offset \u03b5. It is useful to clearly define the various measures of gap that we reference. In numerical results we have defined the gap H previously (Fig. 2c). In the context of experimental results, H is the actual gap of the plates. The apparent gap Ha is that based on instrument zero-gap procedure. The corrected gap Hc is the apparent gap plus a calibrated correction for gap offset \u03b5, Hc \u00bcHa\u00fe\u03b5: \u00f08\u00de The gap offset correction \u03b5 influences calculations of shear-rate, apparent viscosity, and Reynolds number. In terms of the corrected normalized apparent viscosity \u03b7c=\u03b70, \u03b7c \u03b70 \u00bc 2MHc \u03c0 \u03a9R4\u03b70 \u00bc 2M Ha\u00fe\u03b5\u00f0 \u00de \u03c0 \u03a9R4\u03b70 : \u00f09\u00de Combining Eq. (7) and Eq. (9) gives [41] \u03b7c \u03b70 \u00bc \u03b7a \u03b70 1\u00fe \u03b5 Ha : \u00f010\u00de which we will use to calibrate the gap offset error \u03b5 from measurements of apparent viscosity with flat plates",
            " As we shall see, the comparison of experimental and computational results suggests that this is reasonable within our reported accuracy. The 2D numerical results are compared to the 3D rotational triborheometer measurements with the following calculations. The torque M on a flat plate in a parallel rotating disk configuration, with constant viscosity \u03b7 and circular streamlines is, M\u00bc Z R 0 2\u03c0T\u00f0r\u00der2dr\u00bc \u03c0 2 \u03b7\u03a9 H R4 \u00f012\u00de where T\u00f0r\u00de \u00bc TN\u00f0r\u00de \u00bc \u03b7 \u03a9r=H is the local Newtonian shear stress and \u03a9 is rotational velocity of the disk. Our microtextured plates are patterned only in the region RioroRo (Fig. 2). We assume that the \u00bd0;Ri and \u00bdR0;R regions have the standard local shear stress due to laminar Newtonian flow between parallel disks, TN\u00f0r\u00de. Within the RioroRo microtexture region, we represent the modification of shear stress as T\u00f0r\u00de \u00bc TnTN\u00f0r\u00de (based on the definition of Tn in Eq. (4)). This results in the piecewise integral, M\u00bc Z Ri 0 2\u03c0TN\u00f0r\u00der2dr\u00feTn Z Ro Ri 2\u03c0TN\u00f0r\u00der2dr\u00fe Z R Ro 2\u03c0TN\u00f0r\u00der2dr \u00f013\u00de where Ri, Ro and R are defined in Fig. 2a. Tn is the fractional change (reduction) of shear stress in the microtexture region, i.e. the dimensionless average shear stress, defined in Eq. (4), that is a function of \u00f0ReH ;Wn;Dn;Hn\u00de in general. The 2D simulations assume an extruded texture shape, where the linear texture density Wn \u00bcW=L (along a streamline) is equivalent to the areal texture density. The 3D microtextures are discrete cylindrical cavities. Without knowing the extent to which flow is influenced across streamlines between cylinders, we compare results for identical linear density Wn \u00bcW=L between experiments and simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002308_s12239-015-0017-2-Figure24-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002308_s12239-015-0017-2-Figure24-1.png",
        "caption": "Figure 24. Schematic position of drive train assemblies: crank shaft, DMF-primary mass, DMF secondary mass, clutch and gear input shaft.",
        "texts": [
            " Variation of the component mass of inertia One objective of the application of the dual-mass flywheel (DMF) is to eliminate the excessive rotational vibration caused by the irregularity of rotational combustion engine speed (Reik, 1987). To attain a high performance of the vibration damping, the increase in the primary J1 and the secondary mass of inertia J2a is one of the commonly applied solutions. Apart from the body of the secondary mass, other assemblies with a direct contact to the secondary mass like a clutch housing are parts of the mass of inertia J2a. The simplified schematic positions of J1 and J2a are illustrated in Figure 24. Another option to damp the vibration from combustion engine is to increase the mass of inertia of the whole assembly at the gear side without raising the sum of entire mass to be engaged. Practically, the mass of inertia of the clutch disk and the gear input shaft J2b can be increased for this purpose. As an example, this section demonstrates different damping effects of the entire drive train and the consequent comfort evaluations due to the variation of J2a and J2b. The variation of the mass of inertia is carried out on the basis of the original vehicle parameters used in the simulation model: J2a = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000465_s00170-013-4872-6-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000465_s00170-013-4872-6-Figure10-1.png",
        "caption": "Fig. 10 The optimized tool path",
        "texts": [
            " Non-linear errors Et are computed at the densified points. They are compared with a given upper bound E0. If Et>E0, more points on the spline (for regular points) or on the line segments (for feature points) are taken for non-linear error computation and comparison until all the non-linear errors are smaller than the given bound. The densified points are converted into NC codes for machining by the machine kinematics. By non-linear error checking and processing, the final tool path is obtained, as in Fig. 10. After cubic spline interpolation and non-linear error processing, the optimized tool path becomes much smoother, which not only makes the machine tools working more stable but also offers higher accuracy for the machining parts. 3.3 The novel postprocessor Integrating the tool path optimization, a novel postprocessor is proposed. Figure 11 shows a schematic diagram of the proposed postprocessor. By the procedure, after the original cutter location source file (CLSF) is read, the lengths of line segments and the angles between two neighboring line segments are computed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003073_060201-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003073_060201-Figure1-1.png",
        "caption": "Fig. 1. (color online) Domains of physical model.",
        "texts": [
            " The numerical simulation electrode induction melting method used included the following steps: (i) choosing the alternating current and heat transfer module in the software COMSOL Multiphysicsr; (ii) using the 2D axisymmetric method to solve a complex electrode induction melting process. The second step also allowed the investigation of the electromagnetic field near the electrode and thermal field effects on the fluid flow with free surface. Considering the physical model and boundary condition problems, some parameters such as air flow, electrode and induction coil were selected as shown in Fig. 1. Parameters including electrode diameter, electrical conductivity, thermal conductivity, density, heat capacity at constant pressure, relative permeability, induction coil frequency, output power, and the distance between the coil and the electrode were set to be 40 mm, 67 S/m, 14 W/(m \u00b7K), 8030 kg/m3, 413 J/(kg\u00b7K), (400 kHz/40 kHz), (100 kW/120 kW), and 9 mm, respectively. Figure 2(a) shows the initial profile of a superalloy electrode immersed in a conical induction coil. Figure 2(b) shows the melting experiment with a frequency of 400 kHz at 100 kW power"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000747_1077546313483786-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000747_1077546313483786-Figure2-1.png",
        "caption": "Figure 2. Diagram for the kinematic pair with clearance.",
        "texts": [
            " Only the dynamic response of the vehicle due to lateral excitation is considered. 3. The steering wheel is fixed in the position of directing ahead in line. 4. Other clearances in the steering system are neglected. As a result, there are four d.f. in the present model, which include the shimmy angle of the right front wheel 1, the shimmy angle of the left front wheel 2, the rolling angle of the front axle about its longitudinal centerline , and the swing angle of the tie rod . The diagram for the kinematic pair with clearance is shown in Figure 2, where the radius of the shaft pin and shaft sleeve is R1, R2 respectively; the clearance in the revolute pair is r\u00bcR2 R1; the center of the shaft pin is O1, and the center of the shaft sleeve is O2; e is the distance between O1 and O2. The interaction force between the two components of the kinematic pair in contact depends on their surface mechanical properties. It is assumed that the surface of one component is flexible; friction is considered, and the other component is regarded as a rigid body"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003915_j.jmapro.2019.05.002-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003915_j.jmapro.2019.05.002-Figure12-1.png",
        "caption": "Fig. 12. Temperature field isothermal faces for a) single bead sample, b) 2nd pass of 10 bead sample, c) 4th bead of 10 bead sample, and d) 10th bead of 10 bead sample.",
        "texts": [
            " longitudinal and transverse normal stresses, have been determined for the points X1, X2 and X10 (refer to Fig. 2) at different depths. The schematic presentation of the experimental set-up for the conducted ICHD residual stress measurement is shown in Fig. 11. It should be noted that, according to ASTM E837-01 [24], the measured residual stresses by ICHD are reliable up to depth of about 0.4 of the mean diameter of the strain gage circle, D. In our experiment, TML strain gages (FRS-2-11) were used and the centre hole diameter was 1.926mm. So, residual stresses were measured up to 0.8 mm deep. Fig. 12 shows the calculated results for isothermal surfaces during the laser cladding process. Due to local heat absorption of the laser beam, the temperature gradient is high in the vicinity of the laser spot. Fig. 12a shows the temperature field for the first pass of the process. It shows symmetry on laser movement direction. In contrast, in Fig. 12b\u2013d, isothermal curves have tendency toward already existed clad beads. The obtained FE temperature results at points X1, X2, X10, and Y10 (refer to Fig. 2) versus time are shown in Fig. 13. It is clear that the temperature variation in multiple-bead cases is cyclic, particularly for ten-bead case with highly cyclic response. The peak temperature at point X10 (placed near to the start point of the laser path) for the second and the third cycles is larger than the first one. On the other hand, for the point Y10 (placed near to the end point of the laser path), the temperature peak increases as the laser proceeds"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003429_1.5035054-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003429_1.5035054-Figure1-1.png",
        "caption": "FIGURE 1. (a) Tool for bending-process (b) Cladded specimen for bendability tests (c) Cladded specimen for tensile tests",
        "texts": [
            " In a first parameter study the layers were cladded in one step without stops for cooling or external cooling (Batch 1). In a second parameter study (Batch 2) an in-house designed cooling jig was used. To achieve a thickness of 1 mm a second layer was added in Set 2.2. 160028-2 The specimens will be cladded with different layer thicknesses on one side, which will be 25% and 50% of the base layer (2mm), which results in cladding thicknesses of 0,5mm and 1mm. The tensile tests were conducted on a Zwick 1486 universal tensile-testing machine. For the tests, flat specimens according to DIN 50125 were used (see Fig. 1 (a)), which were mounted with manually strained wedge clamp heads. For the manufacturing of the specimens, first rectangular areas are deposited on the base material via LMD. It is important, that the cladding-area is large enough, so that the specimens can be extracted from these cladded areas without including the edges. Finally, the specimens are finished with a milling process to ensure the necessary surface qualities. The bending tests were also conducted on a Zwick 1486 testing machine. In this case, rectangular bending specimens with 90 x 10 mm dimensions were used, see Fig. 1 (b). A special tool was used to pre-bend the specimens, shown in Fig. 1 (c). Afterwards the specimens are bent to 180% between to plane tools. The specimens are cladded on one side, which is also specified in Fig. 1 (c). The criterion for bendability is the angle at which macroscopic cracks occur. Aim of these bending tests is to ensure that the bendability is still given after the cladding process is conducted on the material. 160028-3 RESULTS Manufacturing of cladding A large number of pores - many of them lager than 100 \u00b5m in diameter - was present in the specimens produced in Batch 1 with 0.5 mm and 1.0 mm layer thicknesses, see Fig. 2. This can be attributed to overheating of the melt. Due to the low thickness of the blank heat dissipation from the melt is restricted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002554_1.4033128-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002554_1.4033128-Figure1-1.png",
        "caption": "Fig. 1 Schematic of pad bearing with multiple dimples on pad",
        "texts": [
            " The other mechanism is the enhancement of the entire wedge action over the lubricated area by incorporating textured features. The present authors suggested the mechanism and named the balancing wedge action [2,3]. When a single dimple is created at the inlet or outlet side of the pad with a pivot point, the incorporation of the dimple changes the pressure distribution over the lubricated area to incline the pivoted pad. This process improves pressure generation over the lubricated area [2,3]. In the current study, the authors extend the concept of the balancing wedge action to multiple dimples. Numerical Model. Figure 1 depicts a schematic of an infinite pad bearing. The upper stationary surface can be freely rotated about the pivot point xpv. On the stationary surface, multiple dimples with a height hd and width ldw are created within the dimple portion of width ld, which starts from xdst. Each dimple has a land width llw at the leading side or trailing side. Figure 1 shows a textured pad for the case in which dimples have lands at the leading side. The lower flat surface moves from left to right. The flow of lubricant in the lubricated area is assumed to be incompressible and isothermal. The origin of the coordinate x is located at the inlet. The film thicknesses at the inlet and outlet are defined as h1 and h0. Dimensionless forms of the variables are as follows: Hd \u00bc hd h0 (1) N \u00bc ld ldw \u00fe llw\u00f0 \u00de (2) Ld \u00bc ld l (3) Ldw \u00bc ldw l (4) Llw \u00bc llw l (5) a \u00bc ldw ldw \u00fe llw\u00f0 \u00de \u00bc Ldw Ldw \u00fe Llw\u00f0 \u00de (6) Xpv \u00bc xpv l (7) where a is the dimple width ratio"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001988_3dp.2014.0017-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001988_3dp.2014.0017-Figure3-1.png",
        "caption": "Figure 3. STL model of a turbine as viewed through commercial CAD software ( left ), corresponding 3DP sliced image for a given cross section ( center ), and corresponding SIS sliced image ( right ).",
        "texts": [
            " Open - source and free software were then used to view, manipulate, and heal the STL files when necessary. The layered images were constructed using a predetermined uniform layer thickness of 120 \u03bc m for a given binary STL file format input. A dedicated graphical user interface was developed in Visual C# for STL slicing and image generation in the SIS process. SIS sliced image generation differs significantly from other AM processes because only the periphery of the part is marked for inhibitor deposition. Figure 3 illustrates the difference between a binder jetting (3DP) image and an SIS sliced image. The 3DP sliced image marks the entire part cross section for binder deposition. In contrast, the SIS sliced image marks the outer part periphery only. The inhibitor solution chosen for this research consists of sucrose (C 12 H 22 O 11 ) dissolved in water and an organic surfactant. Previous research by Yoozbashizadeh analyzed the inhibition effects of various ceramic salts and carbohydrates for bronze material"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003797_s00170-018-03239-z-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003797_s00170-018-03239-z-Figure4-1.png",
        "caption": "Fig. 4 Model of skiving tool cutting angles, a illustration of skiving process by a single-blade tool, b mathematical model of cutting angles",
        "texts": [
            " Although the structure of the multiblade skiving tool is similar to Mitsubishi\u2019s MHI superskiving cutter, the detachable subblade of multiblade tool, which makes it convenient to regrinding of the rake face and unnecessary for subblade with high precision, is better than the integral blade of MHI super-skiving cutter. Consequently, the multiblade skiving tool can be expected to realize higher efficiency and longer tool life with the same machining accuracy as the skiving tool with one cutting blade. Figure 4a presents the illustration of skiving process by a single-blade tool, the designed rake angle and relief angle are \u03b3o and \u03b1o, while the tool working rake angle and relief angle are \u03b3oe and \u03b1oe, respectively. In order to obtain the working angles of tool, the geometric model of skiving tool\u2019s cutting angle is established in Fig. 4b. Since the crossed axes angle\u03a3 between workpiece and tool, the generatingmotion of tool is the elliptic equation from the front section view of workpiece. As shown in Fig. 4b, the parametric equation of edge path can be represented as follows: x \u00bc acost y \u00bc bsint \u00f04\u00de where, a \u00bc r f \u00fe \u2211 n i\u00bc1 \u039bi and b = a \u22c5 sin\u03a3, rf is the root radius of tool, \u039bi is the cutting depth per cutting pass, i is the number of cutting pass, t is the parameter of ellipse in Cartesian coordinate. Assuming that one point on the ellipse is k (xk, yk), the horizontal axis coordinate is determined by the following: xk \u00bc a\u2212 \u2211 k i\u00bc1 \u039bi \u00bc r f \u00fe \u2211 n i\u00bck\u00fe1 \u039bi \u00f05\u00de From the parametric Eq. (4), x = a cost; thus, the parameter angle of ellipse at the point k can be derived as follows: tk \u00bc acos xk a \u00f06\u00de One of the cutting angles of tool at the point k can be formulated as follows: \u03b3k \u00bc atan y 0 x0 \u00bc \u2212atan b a cottk \u00f07\u00de where, x 0 \u00bc \u2212asintk y 0 \u00bc bcostk , a \u00bc r f \u00fe \u2211 n i\u00bc1 \u039bi \u00bc mnzt 2cos\u03a3 \u22121:25mn \u00fe \u2211 n i\u00bc1 \u039bi, b = a \u22c5 sin\u03a3, \u039bi is the cutting depth per cutting pass"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001853_s10846-015-0233-z-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001853_s10846-015-0233-z-Figure8-1.png",
        "caption": "Fig. 8 Leg parameters definitions",
        "texts": [
            " \u2013 Determine the necessary displacement to counter the perturbation. \u2013 Compensate the external perturbation by moving the CG of the robot or moving the foot contacts of the robot. The following presents the experimental demonstration of foot force based reactive stability control using a radially symmetric hexapod robot. 4.1 Experimental Hardware Figure 7 shows the radially symmetric hexapod robot used in the experiment. The diameter of the platform is 300 mm. The legs of the robot have three joints as shown in Fig. 8. Each leg includes three separate segments which are connected together by revolute joints. The names, magnitudes, and limitations of the leg segments and joints are listed in Table 1. All of the legs are connected to the main body (platform) of the robot through hip joints. The hexapod robot consists of a Lynxmotion hexapod robot kit [25] and a Gumstix Verdex Pro XM4-BT COM tiny computer [26]. The hexapod robot has 18 HS-485HB servos controlled by a SSC-32 sequencer. There is a built-in proportional controller for each servo"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003356_j.apm.2018.01.018-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003356_j.apm.2018.01.018-Figure6-1.png",
        "caption": "Fig. 6. 3D representation of the expanded 2D tyre model.",
        "texts": [
            " The 3D tyre model was created by expanding the 2D axisymmetric model into different repetition angles using the Axisymmetric Model To 3-D feature in Marc/Mentat. These angles are shown in Table 5 . This expansion was based on mesh sensitivity studies completed by Maritz [5] and Van Blommestein [6] , as well as the work presented in [7\u20139,17,18] . The rubber components and the bead were meshed with Full & Herrmann Formulation linear hexagon elements. The body ply and steel belts were meshed with quadrilateral Membrane Rebar elements. The 3D model consisted of 88560 elements and 99121 nodes. The 3D representation of the expanded 2D tyre model is in Fig. 6 . The results obtained from the loading and rolling analyses were used to obtain its the elastic strain energy density at one tyre revolution. The 2D axisymmetric tyre model were used to simulate the temperature distribution in the tyre crosssection. Both the the 2D and 3D numerical models were modelled as implicit non-linear structural finite element models. The 2D numerical model developed for the inflation analysis is shown in Fig. 7 . The tyre model was inflated using a Cavity Pressure Load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001042_0959651817698350-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001042_0959651817698350-Figure11-1.png",
        "caption": "Figure 11. Manipulator configuration stage I.",
        "texts": [
            " l14 and l47 are link lengths for new configured planar manipulator. The tip velocity can be found by differentiating equations (17) and (18) as _X# tip= _XCM r( _f) sf l14( _f+ _u1 + _u1n) sf11n l47( _f+ _u1 + _u1n + _u4n) sf11n4n \u00f019\u00de _Y # tip= _YCM + r( _f)cf + l14( _f+ _u1 + _u1n)cf11n + l47( _f+ _u1 + _u1n + _u4n)cf11n4n \u00f020\u00de Equations (19) and (20) help in evaluating transformer moduli for drawing bond graph model. The obstacles present in the workspace are recognized by providing a barrier encircling the obstacles as shown in Figure 11. Assume that the obstacle 1 has hexagonal shape and obstacle 2 is of square shape. The mathematical expression for obstacle detection29,30 can be given as djLO= f(xi, yi) \u00f021\u00de where djLO denotes the pseudo-distance between the manipulator link and j number of obstacles and f(xi, yi)= (xi xci) 2 + (yi yci) 2 ai 2 \u00f022\u00de The coordinates xi and yi represent the tip position near obstacles. The coordinates xci and yci are the center of mass coordinates of the obstacles barrier and ai is the radius of the barrier",
            " Then, one of the FJP needs to be made active simultaneously, and hence, the robot configuration needs to be changed accordingly to continue the desired motion planning. At minimum positive value of the pseudo-distance, the robot configuration needs to be transformed. However, as much as a positive value of pseudo-distance can provide safest motion planning, but it may reduce workspace. This concept can be extended for the proposed robot system as follows. Stage I. This stage is schematically represented in Figure 11. The circumstance when both the obstacles collide with the manipulator is shown in Figure 11. The whole manipulator is divided into two sections for simplicity. In the first section of the manipulator, only joint 1 is active, and in the second section, only joint 4 is active. At this stage, d1LO \\ 0 for obstacle 1 is assumed in section 1 and d2LO \\ 0 for obstacle 2 is assumed in section 2 of the manipulator. These mean that the robot system suffers a serious collision from both the obstacles. Now, assume that the planar space robot system avoids both the obstacles one by one. Stage II. This stage is schematically represented in Figure 12",
            "29 From Figure 15(a), it is observed that the tip closely follows the designated path. In Figure 15, 1,4A denotes that joints 1 and 4 are active joints. Figure 15(b) represents the tip trajectory when the manipulator avoids obstacle 1 but collides with obstacle 2 (stage II). This stage can be expressed as partial collision-free trajectory. In Figure 15, part A of trajectory is the resultant of joints 1 and 4 actuation. During tracing of this part of trajectory, the first section of manipulator (Figure 11) arm comes close to obstacle 1. Furthermore, if this motion still continues, this section of manipulator would have collided with obstacle 1. Hence, the motion of this section of manipulator is restricted by making joint 1 as FJP, and it is replaced by joint 6 for further motion. The joint 6 was at FJP but now acting as active joint. Hence, joints 4 and 6 are now active which are denoted by 4,6A in Figure 15(b). From Figure 15(b), it is seen that the robot closely track part A of the designated path, but does not the rest part (part B). This trajectory error occurs due to insufficient torque generation by the active joints 4 and 6. Figure 15(c) shows the tip trajectory when the whole manipulator avoids both the obstacles completely (stage III). This stage can be expressed as complete collisionfree trajectory. During tracing the end of part B, the second section of manipulator (Figure 11) comes close to obstacle 2 as it is detected by position sensor. Hence, joint 4 is now kept at FJP which influences the motion of the second section of manipulator, and this joint is replaced by resuming joint 1 again as active joint for further motion with collision free. And the robot tip closely traces the given reference input. In Figure 15(c), 1,6A represents that joints 1 and 6 are active joints. Figure 16 shows the tip trajectory tracking error between the circular reference and the actual tip position of 6-DOF planar space robot: stage I (Figure 16(a)), stage II (Figure 16(b)) and stage III (Figure 16(c))"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000612_ijsurfse.2014.065817-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000612_ijsurfse.2014.065817-Figure1-1.png",
        "caption": "Figure 1 (a) Schematic representation of smooth porous journal bearing, (b), (c) and (d) surface texture at configurations 1, 2 and 3, respectively",
        "texts": [
            " Lin (1997) investigated the rheological effects of couple stress fluids on the lubrication performance of finite journal bearing and concluded that when compared with the Newtonian-lubricant case, the effects of couple stresses provide an enhancement in the load-carrying capacity, as well as reduction in the attitude angle and friction parameter. A number of researchers have investigated the performance of journal bearings both with and without texturing and also numerically as well as experimentally. They have used various kinds of surface texturing techniques such as laser texturing, chemical etching, focused beam, indentation, etc. and also different kinds of texturing. In this work, we imposed sinusoidal wave type of texture at different locations on the porous bearing surface as shown in Figure 1. Etsion and Burstein (1996) examined the hemispherical pores on a square array for a fixed mechanical seal design. The authors found that asperity size significantly affects friction, leakage and film thickness over a wide range of asperity distributions. Etsion et al. (1999) have developed an analytical model to predict the relation between the opening forces and operating conditions in a laser textured mechanical seal. Ronen et al. (2001) demonstrated a model to study the potential use of micro-surface structure in the form of micro pores to improve tribological properties of reciprocating automotive components",
            " The centre differencing scheme of Finite difference method has been used for computation of fluid pressures and other performance parameters by converting the equation into discretised form. Simpson rule is used for calculating bearing performance characteristics. The continuity and momentum equations governing the motion of the lubricant in the absence of body forces and body couples are given by equations (1) and (2), . 0\u2207 =V (1) 2 4= \u2212\u2207 + \u2207 \u2212 \u2207 DV\u03c1 p \u03bc V \u03b7 V Dt (2) where \u2207p is the pressure force, V is the velocity vector. The journal moves at a surface velocity u and the oil film thickness is given by (1 cos )= +h c \u03b5 \u03b8 (3) Figure 1(a) shows the schematic diagram of a smooth porous journal bearing. Surface texture has imposed on the bearing surface at three different locations for predicting the performance. Figure 1(b) shows the first texture location, i.e., configuration 1 and correspondingly other texture locations, configuration 2 and configuration 3 in Figures 1(c) and 1(d), respectively. The flow through the sinter is governed by Darcy\u2019s law which can be given mathematically as \u2202 = \u2212 \u2202 pv y \u03bc \u03c6 (4) where v is the velocity of flow or flow rate in the porous region, \u03c6 is the permeability of the material and negative sign in the equation (4) shows that the flow is in the direction of decreasing pressure. The Reynolds equation for porous journal bearing can be given below as: 3 3 0 6 2 = \u23a1 \u23a4\u239b \u239e\u2202 \u2202 \u2202 \u2202 \u2202\u239b \u239e \u239b \u239e \u239b \u239e+ = +\u23a2 \u23a5\u239c \u239f\u239c \u239f \u239c \u239f \u239c \u239f\u2202 \u2202 \u2202 \u2202 \u2202\u239d \u23a0 \u239d \u23a0 \u239d \u23a0\u23a2 \u23a5\u239d \u23a0\u23a3 \u23a6y p p dh ph h \u03bc U x x z z dx y \u03bc \u03c6 (5) The sinusoidal wave texture equation is given as: sin( )=s\u03b4 A R (6) * * = \u03c0 r \u03b8R w (7) w is sinusoidal wavelength of asperity and equation (7) is for transverse wavy surface texture"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003797_s00170-018-03239-z-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003797_s00170-018-03239-z-Figure1-1.png",
        "caption": "Fig. 1 Machine setup of power skiving for an internal cylindrical gear",
        "texts": [
            " (MHI) have established a new process technology with MHI super-skiving system using barrel-shaped skiving tool for longer tool life and enhanced efficiency in the manufacturing of internal gears. The aim of this study is to propose a novel skiving tool with multiple subblades on the rake face with respect to the traditional tool. Then, the correlation between the cutting parameters and tool performance will be investigated, and the relation between the tool structure and chip deformation will be analyzed as well. The machine setup of power skiving for an internal cylindrical gear is shown in Fig. 1. Consider that the Cartesian coordinate systems S1(O1-x1,y1,z1) and S2(O2-x2,y2,z2) are rigidly connected to the coordinate of workpiece and skiving tool, respectively. The internal gear rotates about z1-axis with an angular velocity of\u03c91 while the skiving tool rotates about z2-axis with an angular velocity of \u03c92. Both rotations are synchronized by an electronic gearbox. The tool center is oriented out of the workpiece center by a radial distance L, as well as a crossed axes angle \u03a3 between the rotational axis of workpiece and tool"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001985_ecce.2015.7309702-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001985_ecce.2015.7309702-Figure5-1.png",
        "caption": "Fig. 5. End shield structure of bearing and lubricant supply for reconfigured motor with custom-built journal bearing (motor B)",
        "texts": [
            " Open type bearings can be re-greased by injecting grease through the inlet and channel, and the internal bearing cap prevents the grease from entering the motor [32]. The rolling element bearings were removed and replaced with precision-machined steel journal bearings with shaft-bearing clearance of 60, 75, 90 \u03bcm (motors B1-B3), as shown in Figs. 5 and 6(a). The grease inlet/outlet, and channel were used for supplying the lubricant (oil) into the journal bearing through gravity feed, as shown in Fig. 5. It is very important to ensure that the oil is supplied evenly throughout the bearing-shaft contact for maintaining the oil film for stable operation. To direct the oil to the center of the bearing axial length, an adapter was added to the end shield to guide the oil to the top of the journal bearing and shaft contact, as shown in Figs. 5 and 6(b). Holes were drilled in the axial and radial direction in the upper part of the bearings to guide the incoming oil from the adapter to the shaft-bearing clearance, as shown in Fig. 6(a). The path of the oil flow through the inlet, channel, adapter, bearing, shaft, channel, and outlet is shown in Fig 5 and in the bearing design drawing shown in Fig. 6(b). It is also critical to control the \u201cflow\u201d of oil to maintain the oil film in the bearing for operation of the motor within its thermal limits. The flow of ISO viscosity grade 68 oil was controlled by feeding the oil into the grease inlet of the end shield through a valve, as shown in Fig. 7, and the oil leaving the grease outlet was circulated back to the inlet. Rubber seals were placed on the shaft to minimize the leakage of oil. The motor was fixed to a 42"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000405_s00170-015-6830-y-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000405_s00170-015-6830-y-Figure4-1.png",
        "caption": "Fig. 4 FE model for hydro-bending of the bi-layered tube",
        "texts": [
            " Table 1 Mechanical properties of inner and outer tubes Material Yield strength \u03c3s (MPa) Tensile strength \u03c3b (MPa) Total elongation (%) Uniform elongation (%) Hardening exponent Inner tube Aluminum alloy 5A02 80 183 20.1 13.9 0.17 Outer tube Mild steel 266 572 37.0 26.8 0.26 2.4 FE model The finite element analysis code ABAQUS 6.8 was employed to study the hydro-bending of bi-layered tube. In the FE model, only half part of the tube was actually analyzed due to symmetry of the problem, as shown in Fig. 4. The upper die and lower die were assumed rigid and modeled as discrete rigid surfaces, while the mild steel/Al alloy tubes were modeled separately using S4R deformable shell elements, with the element size of 4 and 2 mm, respectively. The Coulomb friction model was used for all contact surfaces with the coefficient of 0.1. A constant pressure was applied to the inside surface of the inner tube and the sealing ends. 3 Results and discussion 3.1 Wrinkling behavior Figure 5 illustrates the wrinkling defect and the axial stress under different levels of internal pressure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003278_0954406217743271-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003278_0954406217743271-Figure8-1.png",
        "caption": "Figure 8. Illustration of the role of the tail force on the torque generation at preparatory stroke (a) and (b) time courses of torque generated by the entire body, the anterior and the posterior part.",
        "texts": [
            "9,10 This observation also shows that the wake capture mechanism, which is commonly used on insect flight, is not significant at the larval fish quick turning. Intuitively, whatever the fish intends to capture a prey or escape from a dangerous circumstance, quick acceleration and fast movement are greatly favoured to reduce the time for moving the intended distance. However, many fish decelerate during the quick start and turn case.3,25 Why do the fish undertake this process? A simple model was proposed to answer this question. Figure 8(a) shows the situation at t=T \u00bc 0:5, when a considerable deceleration takes place. From the pressure distribution in Figure 6(a) and (b), the large pressure is generated at the tail, and the force directs to opposite to the escaping direction. The torque is calculated based on the mathematical formula T \u00bc R r0 f dS, where r0 is the position vector with respect to the instantaneous centre of mass. Under this definition, d dx \u00f0 r0i mivi\u00de \u00bc r0i Fext i is satisfied, where vi and Fext i are the velocity and hydrodynamic force in laboratory coordinate system, respectively. We found that the resultant torque on tail is in the clockwise direction (Figure 8(b)). This indicates that it is the tail force that causes the deceleration, drives the yaw turning of the body to the escaping direction. The time courses of external torque generated on anterior and posterior parts are also shown in Figure 8(b). A large clockwise torque (negative value in the plot) is observed around the time t=T \u00bc 0:5 on the tail. For the real and untouched larval fish, the flapping of pectoral fins during quick turning is observed,15 which can potentially contribute to torque for turning. However, since pectoral fins\u2019 surface area is so small compared with the tail area, their contribution is also much smaller than the tail contribution. In this paper, we used a fluid\u2013body interactionmodel to study the hydrodynamics of quick turning of the larval fish"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003543_s11665-018-3534-0-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003543_s11665-018-3534-0-Figure1-1.png",
        "caption": "Fig. 1 Mould geometry, pattern, and sampling",
        "texts": [
            " All parameters of the experimental studies are summarized in Table 1. Reduced pressure test samples were collected for bifilm index (Ref 45) measurement as an indication of melt cleanli- ness. A vertical sand mold which has four different steps were used to investigate microstructure and mechanical properties. Steps of the mold were in different thickness as 30, 20, 15, and 10 mm. 30 mm step acted as feeder. Other steps were cut to produce tensile test samples and microstructure samples. Cylindrical samples whose dimensions are given in Fig. 1(a) were machined using a CNC device according to ASTM E 80 standards in terms of dimensions of the samples. The mold geometry, pattern, and sampling are shown in Fig. 1. Porosity formations in tensile test samples of the cast parts were characterized as both volumetric and areal measurements. Archimedes method was applied for volumetric porosity measurement, and images of surface of the tensile test samples were analyzed by via image analysis software for areal calculations. Microstructures were examined from the middle of the cast parts, which is shown in Fig. 1(d). SDAS, size, area, and density of Si morphology were measured for characterization of the microstructure. Fracture surfaces of tensile test samples were analyzed by scanning electron microscopy (SEM). Mechanical test results were analyzed statistically by Minitab. Figure 2 shows the change in melt quality depending on holding time from 30 min to 240 min and degassing process. Figure 2(a) shows that as holding time increases, bifilm index increases when the melt is not degassed. Figure 2(b) displays how bifilm number changes with holding time and degassing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000301_s12555-012-0243-6-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000301_s12555-012-0243-6-Figure2-1.png",
        "caption": "Fig. 2. 3-axis FT (Force/Torque) sensor.",
        "texts": [
            " The major strength of the Cartesian computed torque method is to set joint compliances independently in Cartesian coordinates. We can obtain different stiffness and damping characteristics easily on each Cartesian axis independently according to the gain setting. In the case of \u03b84, \u03b85, used to determine the orientation of the end-effector, we simply use PDservo because the orientation of the end-effector is not crucial for compliance. 2.2. 3-axis force/torque sensor A 3-axis FT sensor, shown in Fig. 2, was developed for measuring external force and moments on the endeffector. We can measure the normal force along the Z axis, Fz, and the moments along the X, Y axis, Mx, My. The diameter of the FT sensor is 80mm, and the thickness is 24mm. In the FT sensor, there are 12 strain gauges consisting of three Wheatstone bridges and an FT sensor board that includes an amplification circuit and microprocessor. The FT sensor board is installed inside the FT sensor structure. The only outputs are the amplified and calibrated digital values of the force and torque data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001956_citcon.2015.7122604-Figure16-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001956_citcon.2015.7122604-Figure16-1.png",
        "caption": "Fig. 16: Ceramic coated outer ring",
        "texts": [],
        "surrounding_texts": [
            "Various methods exist for correcting motor and VSD shortcomings, such as correcting the cable shielding, improving the grounding of the motor housing, using single phase filters, installing an electromagnetic shield between the stator and rotor, grounding the rotor, or using a non-conductive coupling. However, one of the simplest methods is to insulate the bearing(s) and disrupt the current flow. This can be done by installing an insulating shield in the bearing housing bore or by insulating the bearing itself. Using the bearing as the insulator is often the easiest and most effective method. Using the bearing as an insulator or disruptor of high frequency current flow can be done in two ways. The first is by adding a ceramic coating to the outer diameter of the outer ring or the bore of the inner ring (Fig. 15). The second is to use rolling elements that are made of ceramic material. Depending on the bearing size, one method may be more economical than the other. The ceramic ring coating is made of aluminum oxide CAlz03) and is nominally 100 microns thick. This coating is a pure insulator in DC motors with a resistance of 50 Mega ohms or more. The dimensions and tolerances of the bearing remain the same as that of a standard bearing without the coating. Figs.I7 & 18. 978-1-4799-5580-0/15/$31.00 \u00a9 IEEE 2014-CIC-1045 In AC motors with high frequencies produced by PWM converters, the ceramic coating or rolling element does not act as a pure insulator or resistor. In this case, an equivalent circuit diagram must be constructed that considers all of the bearing components, the Ah03 coating, and housing. The Al203 coating has to be modelled as a parallel connection between a resistor and capacitor which means we must consider the impedance. The consideration of impedance is fairly complex where the ohmic resistance (R), the capacitance (C), and the frequency if) are included. To increase the impedance of the bearing, the capacitance of the coating should be kept as small as possible. The capacitance depends on the surface area coated, the thickness of the coating, and the coating material. Charts below illustrate the effects of impedance vs. frequency and capacitance vs. frequency Fig. 19. Impedance Z: Magnitude of impedance IZI R IZI 1 z J iz + (2rrfC)Z 1 + jwRC R = Resistance (n) 978-1-4799-5580-0/15/$31.00 \u00a9 IEEE Page 8 of 11 Page 9 of 11 2014-CIC-1045 C = Capacitance (F) f = Frequency (Hz) co = 2n/ cr = dielectric constant for the insulating coating material co = dielectric constant in vacuum A = coated contact area s = thickness of the ceramic coating Il1lIedance 101J01J0 E 10000 -- \ufffd Q. 1000 :0; \ufffd :;; 100 v ------\ufffd 10 \u00a7 FrequencYIHzI Capacitance 4,5OE-09 .---------------------.., 4,00E-09 t=::.:::: :::: :::: :::: ::::::========\ufffd \ufffd 3,5OE-09 j:0; 3,00E-09 +-----------------1 :;; 2,5OE-09 ----------------1 \ufffd 2,00E-09 ----------------1 \ufffd 1,5OE-09 ----------------1 \ufffd 1,ooE-09 +-----------------1 5,ooE-10 ----------------1 0,00800 '-----------------' Frequency IHzl - IZI -c Fig.19: Impedance decreases with frequency increase. Capacitance of the bearing depends on the area coated, thickness of the coating, and the coating material properties. Decreasing the capacitance is related to the coating thickness and the surface area coated. While the outer ring is typically coated, another way to decrease capacitance is to have the inner ring coated instead, depending on the size of the bearing. Use of ceramic rolling elements also lowers capacitance. Bearings built using ceramic rolling elements are typically referred to as \"hybrid bearings\" (Fig.20). In hybrid bearings, the ceramic rolling elements are made of silicon nitride. These silicon nitride rolling elements are electrical insulators. For DC current, the open hybrid bearing, without seals or shields, is a pure insulator in the order of several GQ of resistance. The hybrid bearing becomes a very effective tool for breaking DC current paths through the rolling element-raceway contact. 978-1-4799-5580-0/15/$31.00 \u00a9 IEEE 2014-CIC-1045 Fig.20: Hybrid bearing with silicon nitride rolling elements. With high-frequency currents generated by PWMs, the ceramic rolling elements of a hybrid bearing act as a dielectric between the inner and outer rings. The capacitance of a hybrid bearing is very low (in the order of picoFarads). The relatively large diameter of the rolling elements, the small size of the raceway contacts and the low relative dielectric constant, \u00a3\" of the silicon nitride gives impedances above 1 kQ, even at frequencies of 1 MHz. Thus, the hybrid bearing becomes an extremely effective tool to prevent electrical damage to bearings, particularly in respect to high-frequency currents. The lubricant recommended for bearings in electric motors with silicon nitride rolling elements is a polyurea base thickener with an esther oil. Greases with these constituents have been found to have excellent life at high operating temperatures and high speeds found in motors. Grease life with hybrid bearings can be 4 times as long as with standard bearings. The capacitance of either the ceramic coating on the bearing ring or the ceramic rolling elements is typically lower than that of typical insulating sleeves used in the housing and also easier to install."
        ]
    },
    {
        "image_filename": "designv11_13_0003114_lra.2017.2728200-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003114_lra.2017.2728200-Figure10-1.png",
        "caption": "Fig. 10: Crossing over rubble-strewn ground (from [23])",
        "texts": [
            " This took approximately 180s to generate a single valid sample, illustrating this approach is not feasible for planning with contacts as the probability of selecting valid contact configurations is small. 2377-3766 (c) 2017 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Next we ran the RB-PRM planner from the open-source Humanoid Path Planner package [23] on the same robot for two scenarios, the rubble (Fig. 10) and the truss (Fig. 1). The crux of the approach [4] is that if each robot leg is operating within its reachability limits, then it is assumed a stable fullbody configuration can be generated. For the rubble scenario shown in Fig. 10, our planner took on average 44s to generate a motion plan, while the RB-PRM planner took between 10s and 15s which was comparable to the 7s reported in [4] using the HyQ platform. We then tested both approaches on the truss shown in Fig. 1. Our planner successfully generated paths in 77s, but RB-PRM failed to plan the vertical transitions, likely as the reachability assumption did not hold. These results show that on planar terrains our planner is within the same order of magnitude performance as [4], while enabling planning in complex environments such as Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000053_1464419314566086-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000053_1464419314566086-Figure4-1.png",
        "caption": "Figure 4. A staggered SRB.",
        "texts": [
            " Fx1 \u00fe Fx2 \u00bc 0 \u00f024\u00de where Fr1 \u00bc X j\u00bc l j\u00bc0 Q1 j cos cos\u00f0 1j \u00fe \u00de; Fr2 \u00bc X j\u00bc l j\u00bc0 Q2 j cos cos\u00f0 2j \u00fe \u00de Fx1 \u00bc Xj\u00bcZ j\u00bc1 Q1 j cos sin\u00f0 1j \u00fe \u00de; Fx2 \u00bc Xj\u00bcZ j\u00bc1 Q2 j cos sin\u00f0 2j \u00fe \u00de at Kungl Tekniska Hogskolan / Royal Institute of Technology on September 10, 2015pik.sagepub.comDownloaded from In the above equations, is the azimuth angle of No. 1 roller. The induced roller thrust loads in z direction are in the state of self-balance.33 Elastic deformation analysis of roller raceway with no roller error When the rollers are symmetrically distributed in the range of load zone (shown in Figure 4), the SRB inner ring will move along the loading direction under an applied radial load Fr. When the rollers are not symmetrically distributed in the range of load zone (shown in Figure 5), instead of moving along the loading direction, the inner ring will move to a direction that has an angle of \u2019 relative to the loading direction. The value of \u2019 can be determined by simultaneously solving the static equilibrium equations both in x direction and in y direction. According to the deformation compatibility relationship of roller raceway, the normal elastic deformation of roller raceway 1 j and 2 j can be expressed as 1 j \u00bc r cos\u00f0 1j \u00fe \u2019\u00de cos Pd 2 \u00f025\u00de 2 j \u00bc r cos\u00f0 2j \u00fe \u2019\u00de cos Pd 2 \u00f026\u00de where, 1j and 2j are the jth roller azimuth angle in the first and second row respectively 1j \u00bc 2 Z \u00f0 j 1\u00de \u00f027\u00de 2j \u00bc 0 \u00fe 2 Z \u00f0 j 1\u00de \u00f028\u00de Z is the number of rollers for each row, 0 is the staggering angle between No. 1 roller in the first row and No. 1 roller in the second row (as shown in Figure 4). For a aligned SRB, 0 \u00bc 0 or 0 \u00bc 2 =Z, and for a staggered SRB, 0 \u00bc =Z. Substituting equations (25) and (26) into the static equilibrium equations (23) and (24), two nonlinear equations with the two unknowns r and \u2019 can be obtained, which can be solved with the Newton\u2013 Raphson iterative method.34 The roller load value and axis position of the SRB can be calculated when the rollers are in any position by changing the value of . Elastic deformation analysis of roller raceway with roller error In actual SRBs, the rollers generally have some diameter error, which will affect roller load distribution and bearing axis orbit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001698_j.mechmachtheory.2018.12.006-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001698_j.mechmachtheory.2018.12.006-Figure4-1.png",
        "caption": "Fig. 4. Main dimensions of the test vehicle (dimensions in mm).",
        "texts": [
            " For both these conditions, given the speed of the engine shaft and the outlet shaft, the following step are applied: - compute the actual torque on the wheels, given by the most stringent between the maximum pulling torque and the maximum power conditions; - using the planetary gear equilibrium equations compute the working condition of the CVT shaft \u03c9 CVT and T CVT ; - compute both the speed and the torque at the second unit using the value of \u03c4 II obtained in Eq. (14) ( \u03c9 II = \u03c4II \u03c9 CV T and T II = T CV T / \u03c4II , respectively); - obtain the flow rate in the second hydraulic unit, assuming p = p max : Q II = T II | \u03c9 II | p max (17) A 75 kW forklift was selected as a test vehicle. Its main dimensions and design data are reported in Fig. 4 and Table 3 , respectively. The design speed value and the maximum working pressure for the two hydraulic units are respectively \u03c9 Hy Max = 30 0 0 rpm and P max = 400 bar. The results of the preliminary design can be easily rationalized for the considered test case in terms of the two design variables: \u03b8 and v fmp . In order to obtain a continuous map, the parameter \u03b8 was sampled in the whole range 0 . . . 2 \u03c0 , corresponding to the ideal values of the standing gear ratio between \u22121 \u2026\u2212\u221e , for all the six configurations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000210_s11043-015-9260-1-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000210_s11043-015-9260-1-Figure3-1.png",
        "caption": "Fig. 3 CETR UMT3 tribometer fitted with a 12-mm rigid indenter",
        "texts": [
            " The joint capsule is open for approximately 10 min during the process. Previous research highlights the importance of minimizing the exposure time of cartilage to open-air (Smyth et al. 2014). The cartilage plugs are placed in a UMT CETR tribometer. The tribometer imposes a nearly instantaneous (within approximately 30 ms) displacement on the cartilage surface, while tracking the force generated in the cartilage matrix. By design, this is a stressrelaxation experiment. The tribometer holds a 12-mm rigid aluminum cylinder attached to a load cell, as shown in Fig. 3. Initially, the cylinder contacts the cartilage surface with a preload of 0.5 N. The preload ensures that the cylinder makes complete contact with the cartilage surface. In effect, the preload is flattening out any curvature in the cartilage. At time t = 0, a practically instantaneous downward displacement is imposed on the cartilage, and the resulting force is measured. After approximately 180 s of measurement, the rigid cylinder is withdrawn from the surface. The cartilage is allowed two minutes to recover between tests, and the procedure is repeated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002799_978-3-319-42255-8_12-Figure12.3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002799_978-3-319-42255-8_12-Figure12.3-1.png",
        "caption": "Fig. 12.3 Scheme of the measurement area, the solid arrow indicates the direction of the component of displacement measured with the ESPI set-up",
        "texts": [
            " The temperatures comparison between Fig. 12.2b, d shows that increasing the drilling speed from 5000 to 50,000 rpm leads to a remarkable increase of temperature. Indeed, the maximum temperature for 5000 rpm is 31.9 C whereas for 50,000 rpm is 77.9 C. This could explain the difference in the hole quality between the two drilling speeds. In Fig. 12.2a the hole bottom is flat and regular whereas in Fig. 12.2c there is a central bulge that is probably due to a local boiling of the ABS material during the drilling process. In Fig. 12.3 a scheme of the measurement area has been reported. In order to avoid local effect (e.g. plasticization or chips presence) the area considered for the measurements goes from two times to four times the hole radius. In this circular crown, a 7 7 pixels area has been selected to record the displacements after the drilling phase. In this paper, in order to simplify the results only the displacements every five step of drilling, i.e. 0.1 mm of depth, have been reported. Moreover, these values have been averaged in order to obtain one value for the whole area"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001875_20140313-3-in-3024.00016-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001875_20140313-3-in-3024.00016-Figure4-1.png",
        "caption": "Figure 4. Roll/Pitch axis model",
        "texts": [
            " To make a roll angle \u03c6 along the x-axis of the body frame, a differential angular speed is given to rotors \u201e3\u201f and \u201e4\u201f while keeping the whole thrust constant. Likewise, a differential angular speed is given to rotors \u201e1\u201f and \u201e2\u201f to produce a pitch angle \u03b8 along the y-axis of the body frame. The quadcopter is assumed to be symmetric with respect to the x and y axes so that the centre of gravity is located at the centre of the quadcopter and each rotor is located at the end of bars. Assuming that rotations about the x and y axes are decoupled, the motion in roll/pitch axis can be modelled as shown in Fig. 4. As illustrated in the Fig. 4, two propellers contribute to the motion in each axis. The thrust generated by each motor can be calculated from (1) and used as corresponding input. The rotation around the center of gravity is produced by the difference in the generated thrusts. The roll/pitch angle can be formulated using the following dynamics:   FLJ  (2) where, yx JJJ  (3) are the rotational inertia of the device in roll and pitch axes. L is the distance between the propeller shaft and the center of gravity, and 21 FFF  (4) Where F1 and F2 represent the thrust generated by rotors \u201e1\u201f and \u201e2\u201f respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002052_isgteurope.2014.7028952-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002052_isgteurope.2014.7028952-Figure4-1.png",
        "caption": "Figure 4. Cross-section of ACSS/TW conductor.",
        "texts": [
            " [28] INVAR conductors are characterized by having a transition temperature in the region of 85-100\u00b0C, after which all the mechanical strength of the conductor is provided by the INVAR core. Above this transition temperature the sag inhibition effect takes place, so that it increases very little with temperature. Nevertheless, this characteristic means that the performance of ZTACIR conductors is very similar to the performance of ACSR conductors up to the transition temperature. The advantages of ZTACIR conductors become obvious when the operating temperature is about 100-120 \u00b0C. Aluminum Conductor Steel Supported (ACSS) is described in [29]. \u201cFig. 4\u201d shows the cross-section of a ACSS/TW conductor. ACSS consists of fully annealed strands of aluminum (1350-0) concentric-lay-stranded around a stranded steel core. The coated steel core wires may either be aluminized, galvanized, zinc-5% aluminum Mischmetal coated or aluminum clad. The steel core is available in either standard strength or high strength steel. ACSS are in different designs: \u201cStandard Round Strand ACSS\u201d, or \u201cTrapezoidal Aluminum Wire\u201d. A main modification with respect to the traditional ACSR is that the aluminum is totally annealed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002296_s12239-014-0093-8-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002296_s12239-014-0093-8-Figure3-1.png",
        "caption": "Figure 3. Images MPC of the joints H (a) and E (b) from Figure 2.",
        "texts": [
            " (1) The geometry of the entire system including the AB ABi \u2206+( )2 \u21262+= \u2206 m cos\u03b8= \u2126 m 1 \u03b8sin\u2013( )= \u03b8 arc cosn 1 \u03b4cos\u2013( ) m -------------------------= \u03b4 2 arc tan 4 n2 m2 u4 4u3m 4u2m2++\u2013\u22c5 \u22c5 4 n2\u22c5 u2 2mu\u2013+ --------------------------------------------------------------------- + + 2nu 2mn\u2013 4 n2 u2 2mu\u2013+\u22c5 -----------------------------------\u239d \u23a0 \u239c \u239f \u239c \u239f \u239c \u239f \u239b \u239e \u22c5= suspension mechanism and the wheel was created in Autodesk Inventor Professional 2008. (2) The geometry of the mobile frame 2 and upper 3 and lower 6 arms were imported in MSC Patran 2007 as STEP files. The mesh elements Tria 4 have been used. A multi point constrain (MPC) was used to constrain the nodes of the interior cylinder of bushing with respect to the node from the joint center to avoid the application of the loads (forces, moments) in a single node. Figure 3 illustrates two constraint states in MPC images of the mobile frame, for two different types of joints. (3) Patran as preprocessor gives a command to MD R2 Nastran that generates the MNF files for flexible elements (mobile frame 2 and upper 3 and lower 6 arms). (4) The STEP files for rigid components (others the ones deformable) and the MNF files for flexible elements are imported in ADAMS MD R2/View and jointed with rigid or compliant bushings. Then the motions of mechanism components are simulated considering the variable distance AB of the actuator depending on the bump or rebound (applied by a vertical oscillator, Figure 1) to the wheel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure5-1.png",
        "caption": "Figure 5: z-displacement of the periodic lattice structure",
        "texts": [
            " Several software tools for the generation of such geometries for additive layer manufacturing are already available. Structure Geometry. The structure is created by joining one and the same cubic elementary cell in the three directions in space. As mentioned before, a strut diameter of 2 mm and an interval of 5 mm have been chosen. The explained build-up has been applied on the design space in figure 3 (see figure 4). This leads to a total mass of 147.5 g for the whole structure. The bearing and the force application are realized as described before. Stiffness. Figure 5 shows the simulation result for the z-displacement of the structure in figure 4. The displacement in the point of force application, which is the reference point to compare the different structures, is 0.142 mm. Together with the applied force of 300 N, a stiffness of 2117 N/mm results for the structure. To make the performance more comparable to other structures, the ratio of stiffness to mass is calculated. Here, a value of 14.4 N/(mm*g) arises. Maximum Stress. In figure 6 a), the von-Mises comparison stress is depicted for the regular lattice structure under the given load",
            " In order to keep the results comparable, the struts\u2019 diameters have been chosen to 2 mm as in case of the periodic structure. So, the course of the structure is the only differing variable in contrast to the periodic structure. The resulting structure has a mass of 83.57 g, which is 43 % less than in case of the periodic structure. Stiffness. Figure 9 shows the simulation result for the z-displacement of the structure. In order to make the results more comparable, the colour scale is the same as in figure 5. The displacement in the point of force application is 0.139 mm, which means a slight improvement compared to the periodic structure. Together with the applied force of 300 N, a stiffness of 2166 N/mm results for the structure. Due to the extensive reduction of the structure\u2019s mass, the ratio of stiffness to mass improves for 81 % towards a value of 25.9 N/(mm*g). Maximum Stress. In figure 10, the von-Mises comparison stress and the stress in xx-direction in cutting plane 2 are depicted for the flux of force adapted lattice structure",
            " To obtain the described geometry with straight struts, the nodal points of the previous structure have been retained and the curved structure has been replaced by straight beam elements. Analogue to the investigations presented before, the diameters of the struts were set to 2 mm. The resulting structure and its constraints and loads can be seen in figure 12. The mass of this geometry is 83.58 g. Thus, the weight is similar to the curved structure and 43 % below the periodic one. Stiffness. Figure 13 shows the simulation result for the z-displacement of the structure. In order to make the results more comparable, the colour scale is the same as in figure 5 and 9. The displacement in the point of force application is 0.130 mm. Together with the applied force of 300 N, a stiffness of 2317 N/mm results for the structure. This means an improvement of 7 % compared to the curved structure and 9 % compared to the periodic geometry. This leads to a value of 27.7 N/(mm*g) for the stiffness to mass ratio. Hence, an additional increase of 7 % results in contrast to the curved structure. Maximum Stress. Way more important than the enhancement of the stiffness is a decrease of the maximum of the appearing von-Mises stress compared to the curved structure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001585_aim.2018.8452392-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001585_aim.2018.8452392-Figure1-1.png",
        "caption": "Figure 1. PEW-RO V",
        "texts": [
            " However, in order for the robot to output this traction force, a certain size is required for the robot. Therefore, this robot cannot be used in a pipe of diameter 150 mm or less. The above facts suggest that it is difficult to inspect a long distance inside a pipe of diameter 150 mm or less at present. We focused on the peristaltic movements of earthworms that can travel even in a narrow place. We developed a robot with a peristaltic motion that can travel over 100 m in the pumping pipe by reproducing this movement using an axial fiber reinforced artificial muscle (Fig.1) [11]. The target pipe is a 100A pipe (inner diameter 108 mm).When many curved pipes exist within a larger pipe, it becomes impossible for the robot to travel due to insufficient pulling force. In previous research, when the robot passed through the bent pipe, we studied ways to lower the friction by hooking the friction reduction unit to the corner [12]. However, the friction reduction unit does not always operate. Robots that propel using vibration are also being developed [13]. The entire body of this robot is covered with fine hair"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000405_s00170-015-6830-y-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000405_s00170-015-6830-y-Figure8-1.png",
        "caption": "Fig. 8 Non-circularity under different internal pressures. a Non-circularity curve. b Section ovalization of inner tube. c Section deformation of bilayered tubes",
        "texts": [
            " For the bi-layered tubes used in this experiment, the critical value of internal pressure is 0.4py. 3.2 Effect of internal pressure on non-circularity of cross sections Diameters of final tubes were measured to calculate the noncircularity by Eq. (2) for describing the ovalization of cross sections. Values of initial non-circularity listed in Table 2 were considered to obtain more exact values of the non-circularity after hydro-bending. The measurements involve only with the tube specimen without wrinkles as the pressure is larger than the critical value of 10 MPa (0.4py). Figure 8 illustrates the measured non-circularity and section shape of the final inner tube under different levels of internal pressure. It is shown that the distribution of noncircularity under various pressures is like a mountain peak shape; the maximum value of non-circularity is located at the symmetrical section and decreases remarkably with pressure increasing. At the pressure of 10MPa, the non-circularity value is up to 22 %, and there is a flatting portion at the top of the tube. However, the internal pressure reaches 25 MPa; the non-circularity value drops down to 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002445_ijmmme.2016010103-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002445_ijmmme.2016010103-Figure1-1.png",
        "caption": "Figure 1. Transmission gearbox full assembly",
        "texts": [
            " Noise, vibration and harness (NVH) with thermo-mechanical performance are the two important design parameters of compact multi speed transmission. Transmission gearbox consists of arrangement of gears and gearbox casing. The designing and assembly were done using solid Edge (Solid Edge, 2006) and Pro-E (Pro-E, 2013). All designing parameters were obtained from measurements and drawing sheets. The casing encloses the gear assembly, input, output and lay shaft. In this study only direct drive 4-speed gears were considered. Figure 1 shows the full assembly of transmission gearbox of vehicle. The isometric view of vehicle shows casing designed in three parts. The main part covers the gear assembly and other two parts used to enclose the transmission casing. Figure 2 shows 4-speed gear assembly and shafts for deformation analysis. The numerical simulation was performed for the transient structural, thermal analysis of gear train under the influence of load, rotational speed and convection heat transfer coefficient. Gear oil works as gearbox lubricant to cool the gears and transfer the heat through convective process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002867_arso.2016.7736297-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002867_arso.2016.7736297-Figure10-1.png",
        "caption": "Figure 10. The polishing effect of different method",
        "texts": [
            " In accordance with the previous set, took \u0394=0.02mm, \u03b4 =5mm, and the blade could be planned for the 9 horizontal trajectory, a total of 850 robot processing points. The flexible robotic polishing system was set up as shown in Figure 9, the parameters of various devices used in this experiment were in the Table II: Took the robot files generated by EF-Robot Studio into the robot operation system, and verified the results of the experiment. After the blade was polished, the effect was as shown in Fig.10 following. Comparing different polishing schemes, the roughness, waviness and other indicators of proposed method have greatly improved relative to conventional manual teaching method. Roughness Ra in 0.8, compared with the traditional method of Ra1.6 has much significant improvement, and the preparation time of the robot program could be reduced by 80%, greatly improving the processing efficiency. The polishing path planning was closely related to the final surface quality, in the complex curved surface (especially the freeform surface) polishing process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003112_acc.2017.7963101-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003112_acc.2017.7963101-Figure1-1.png",
        "caption": "Fig. 1: Tilt-rotor tricopter configuration",
        "texts": [
            " Unlike a conventional quadrotor, with four propellers, the TRT has three rotors, one of which has the ability to rotate about the longitudinal axis of the aircraft. This configuration saves weight and energy, but poses new control challenges. To address these challenges, we apply retrospective cost adaptive control (RCAC) by \u201cflying\u201d the aircraft in simulation to determine the essential modeling information as well as the resulting performance. RCAC is developed in [4], [5], an overview is given in [6], and application to the NASA GTM model is considered in [7]\u2013[9]. In the tilt-rotor tricopter (TRT), shown in Fig. 1, two propellers rotate in opposite directions compensating the torque, and the third propeller is tilted by a servo motor in order to compensate for the adverse yaw. A. Ansari and D. S. Bernstein are with the Aerospace Eng. Dept., Univ. Michigan, Ann Arbor, MI. ansahmad@umich.edu. A. Prach is with the National University of Singapore. annaprach@me.com The translational and rotational equations of motion of the TRT, derived in the body frame under a rigid body assumption, neglecting the gyroscopic moments due to the rotors\u2019 inertia, drag forces, and moments, and induced pitching moment by the tilted rotor, are given by [10] u\u0307 = rv \u2212 qw \u2212 g sin \u03b8 + Fx m , (1) v\u0307 = \u2212ru+ pw + g cos \u03b8 sin\u03c6+ Fy m , (2) w\u0307 = qu\u2212 pv + g cos \u03b8 cos\u03c6+ Fz m , (3) p\u0307 = Iyy \u2212 Izz Ixx qr + Mx Ixx , (4) q\u0307 = Izz \u2212 Ixx Iyy pr + My Iyy , (5) r\u0307 = Iyy \u2212 Ixx Izz pq + Mz Izz , (6) where the 3-2-1 Euler angles \u03c6, \u03b8, \u03c8 define the roll, pitch, and yaw, and Fx, Fy, Fz and Mx,My,Mz are the components of the aerodynamic force and moment generated by the rotors in the x, y, and z-body directions",
            " The components of the aerodynamic force are given by Fx = 0, Fy = F1 sin\u00b5 = KF\u21262 1 sin\u00b5, (9) Fz = \u2212(F1 cos\u00b5+ F2 + F3) = \u2212KF(\u21262 1 cos\u00b5+ \u21262 2 + \u21262 3). 978-1-5090-5992-8/$31.00 \u00a92017 AACC 1109 Assuming clockwise rotation for the right and tail rotors, and counterclockwise for the left rotor, the aerodynamic moment components are given by [3] Mx = \u2212l3(F2 \u2212 F3) = \u2212l3KF(\u21262 2 \u2212 \u21262 3), (10) My = \u2212l2(F2 + F3) + l1F1 cos\u00b5 = \u2212l2KF(\u21262 2 + \u21262 3) + l1KF\u21262 1 cos\u00b5, (11) Mz = l1F1 sin\u00b5\u2212M1 cos\u00b5+M2 \u2212M3 (12) = l1KF\u21262 1 sin\u00b5\u2212KM\u21262 1 cos\u00b5+KM\u21262 2 \u2212KM\u21262 3, where the distances l1, l2, l3 are shown in Fig. 1. The parameters of the tilt-rotor tricopter considered in this study are given in Table I [3]. In the trim analysis, we consider the hover condition. By equating the total force and moment to zero, we establish analytical expressions for the corresponding control inputs and states. The total force Ftotal acting on the TRT is Ftotal = [Ftotalx Ftotaly Ftotalz ] T, (13) where Ftotalx = \u2212mg sin \u03b8, (14) Ftotaly = mg sin\u03c6 cos \u03b8 +KF\u21262 1 sin\u00b5, (15) Ftotalz = mg cos\u03c6 cos \u03b8 \u2212KF(\u21262 1 cos\u00b5+ \u21262 2 + \u21262 3). (16) The total moment Mtotal is given by Mtotal = [Mx My Mz] T, (17) where Mx, My, Mz are defined by (10), (11), (12). In hover, the gravitational force is compensated by the vertical component of the combined thrust produced by all three rotors. The reaction torques on the TRT generated by the left and right rotors are equal and opposite, and thus cancel each other. The reaction torque on the TRT produced by the tail rotor is compensated by tilting the tail rotor about the longitudinal axis by the angle \u00b5 as shown in Fig. 1. However, the nonzero tilt angle leads to a nonzero side force. The force is compensated by a nonzero roll angle, which, due to the left and right rotors, produces a horizontal force, which, in turn, requires compensation by the tilt rotor. In hover, the translational and rotational velocities are equal to zero, that is, [u v w]Ttrim = 0 and [p q r]Ttrim = 0. Then, equating the left hand sides of the total force (14)\u2013(16) and total moment equations (10)\u2013(12) to zero, and solving for the unknown inputs and states, yields the pitch and roll trim angles and trim controls \u03c6trim = tan\u22121 [ \u2212 l2KM l1(l1 + l2)KF ] , \u03b8trim = 0, (18) \u00b5trim = tan\u22121 [ KM l1KF ] , (19) \u21261,trim = \u221a l2gm (l1 + l2)KF cos\u03c6trim cos\u00b5trim , (20) \u21262,trim = \u221a l1gm 2(l1 + l2)KF cos\u03c6trim, \u21263,trim = \u21262trim "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000249_elektro.2012.6225639-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000249_elektro.2012.6225639-Figure2-1.png",
        "caption": "Fig. 2. FEM analysis of the SRM, flux lines of aligned rotor position, b) unaligned rot",
        "texts": [
            "00 \u00a92012 IEEE 206 model configurations (changes of the m or other parameters) during the executio The static characteristics of SRM ha by means of FEM under four differ conditions. Namely 0%, 20%, 50%, 70% has been in the short circuit that corresp 6 winding turns respectively. The accuracy of the result depends o mesh and accuracy of the input paramet 10.083 nodes have been used. The calcu out for each individual rotor position static condition. The rotor position \u03d1 aligned \u03d1a to unaligned position \u03d1u wit each position the current was changed w range from 1 to 28 A. In the Fig. 2 t magnetic flux lines of health SRM unaligned position can be seen. In co Fig.3, there is the distribution of magn fault SRM for aligned and unaligned p coil \"B\" turns are in short circuit. Magnetic flux linkage calculation The first parameter which has been a linkage versus phase current for differen =f(I,x). The area bounded by maximal by both \u03c8-I curves for aligned and una equal to mechanical energy, which electromagnetic force [1]. In the Fig. 4 curves obtained by means of FEM f positions, if the phase current has been Amps (45o is equal to unaligned rotor p equal to aligned rotor position)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002737_9781118899076-Figure6.6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002737_9781118899076-Figure6.6-1.png",
        "caption": "FIGURE 6.6 Simplified scheme of the thylakoid membrane with indicated proton flows. PSII, photosystem II; OEC, oxygen-evolving complex; PQ, plastoquinone pool.",
        "texts": [
            " This number looks impressive, but, in fact, owing to a relatively small volume of the lumen, such a dramatic change in the H+ concentration can be reached by a transfer of just few protons into the luminal space. On the other hand, due to the buffering capacity of the lumenal surface of the transmembrane pigment\u2013protein antenna complexes, a high number of protons, which appear due to the water-splitting reaction and operation of the PQ shuttle, can be accommodated within the thylakoid membrane interior and drive photosynthetic ATP production. In this sense, the thylakoid membrane can be compared to a capacitor charged by operation of the light-driven electric charge separation (see Fig. 6.6). A total of 8 light photons (4 in PSII and 4 in PSI) are required to transfer 4 electrons from 2 molecules of water to 2 molecules of NADPH. At the same time, the proton gradient generated can give rise to synthesis of at least 2 molecules of ATP. 127PHOTOSYNTHETIC ELECTRON AND PROTON TRANSPORT Similarly as NADPH, ATP can be utilized not only to power the photosynthetic CO2 fixation but also to provide energy for numerous other biochemical reactions. The effective energetic yield of photosynthesis is not considered to be much higher than 30%, also owing to the absorption of more energetic shorter wavelengths than those directly utilized to drive photosynthetic charge separation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003076_tmag.2017.2708748-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003076_tmag.2017.2708748-Figure2-1.png",
        "caption": "Fig. 2. No-load flux line distribution in 2-D cross-section plane of LOMs. (a) SMTLOM. (b) SMTHLOM.",
        "texts": [
            " 1 (b) and (d) show the 3-D structure and its crosssectional view of the proposed SMTHLOM, respectively. With its stator and mover similar to those of SMTLOM, P 0018-9464 (c) 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. SMTHLOM contains back irons and radially-magnetized ferrite magnets to form an additional flux path (hybrid-fluxpath) in the stator outer region. As shown in Fig. 2, with the introduction of additional flux path, the magnetomotive force (MMF) of NdFeB corresponding to the leakage flux is counterbalanced by that of ferrite magnets, and thus the flux leakage in the stator outer region could be reduced so as to improve significantly the utilization of NdFeB magnets. B. Operational Principle Same as SMTLOM, SMTHLOM moves axially, i.e. along the z axis, which is perpendicular to the radial main flux path in the x-y plane (transverse flux). When excited by a dc current, the armature winding produces the equal armature MMF in the two stator poles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002749_j.ijnonlinmec.2016.08.007-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002749_j.ijnonlinmec.2016.08.007-Figure4-1.png",
        "caption": "Fig. 4. Geometry of sphere to conical frustum contact.",
        "texts": [
            " x y z 2 2 2 2 2 2 Then, the minimum of the above function is found as a solution of the following equations \u03bb \u2202 \u2202 = \u2202 \u2202 = \u2202 \u2202 = \u2202 \u2202 =L x L y L z L 0, 0, 0, 0. The resulting set of nonlinear algebraic equations is solved by using the Newton\u2013Raphson method. The sphere intersects the ellipsoid if point r is contained within the sphere ||r p||r R. Then, the contact point is r, while the penetration depth is given by d \u00bc R ||r p||. Finally, the common normal vector n passes from the contact point r and the sphere's center p. This case leads to an axisymmetric problem with respect to the frustum axis, as shown in Fig. 4. By considering the plane defined by the frustum axis and the sphere center, the problem becomes equivalent to determining the intersection between a trapezoid and a circle. The origin of the coordinate system is placed at the center of the bottom face of the frustum. For testing the intersection, a point x in the trapezoid which is closest to the sphere is located. The space outside the trapezoid is divided into five sections and depending on which section the sphere center p belongs Please cite this article as: A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure8-1.png",
        "caption": "Fig. 8. Gear mesh relation of a planetary gear train [29] .",
        "texts": [
            " (34) Two geometric relations of gear meshing are essential for loaded tooth contact analysis of planetary gear sets, namely the tooth clearances between engaged flanks caused by manufacturing/assembly errors and the contact position of each tooth pair. The clearances are important for load analysis because they affect the shared loads of the contact tooth pairs, while the contact positions of tooth pairs determine the meshing stiffness of the gear pairs. In order to simplify the meshing analysis, a planetary gear set with two planet (index 1 and i ) is considered, Fig. 8 [29] . The geometrical relations of these two sun-planet-annulus gear pairs with a separation angel \u03b3 are at first derived. These relations can be later expanded for the other planet gears with a different separation angel \u03b3 . It is assumed here that the planetary gear set has planar and time-invariant errors, e.g. the position errors of pin-holes and the tooth thickness errors. In order to analyze the tooth clearance due to the errors, the planets rotate so that no tooth clearances between all the contact tooth pairs of the gear pairs except the sun-planet i gear pair are present",
            ", interference between the engaged teeth occurs. In order to avoid this interference, the sun gear must rotate reversely with the interference angle. The final clearance among the engaged tooth pair of all the planets and the sun gear can be expressed as \u03b4res ,i = \u03b4i \u2212 min { \u03b41 , \u03b42 , \u00b7 \u00b7 \u00b7 , \u03b4n } , i = 1 to n (36) With the contact position \u03be S1 of the sun gear engaged with the reference planet 1, the contact positions of the other tooth pairs, e.g., \u03be PS1 , \u03be PA1 etc., can be determined from the geometric relation in Fig. 8 based on the involute geometry accordingly. The related equations are already proposed in the previous work [29] , and will not be listed in this paper. The contact position of each tooth pair will change with the rotation angle \u03d5S of the sun gear or the rotation angle \u03d5C of the carrier. Without considering the time-variant errors, e.g., eccentric errors, the unloaded transmission error does not occur. The new contact position \u03be S k of the sun gear can be also determined by the relation [29] , \u03beS k = mod [ \u03c6C \u00b7 | z A | z P + \u03beS1 , \u03c4S ] . (37) With this angular variable \u03be S k the new geometric relation of the tooth pairs is still similar to Fig. 8 , and the relations are also valid for the case. An example from industrial application is used to demonstrate the feasibility of the proposed analysis model. The distributed contact stress on flanks and the load sharing among planets, obtained from the proposed LTCA model, are at first compared with those from the FEM software (MSC.Marc). The effects of the deformations of the sun gear, the carrier with the planet shafts on the distribution of the contact stress are then identified using the LTCA approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002146_j.mechmachtheory.2014.04.012-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002146_j.mechmachtheory.2014.04.012-Figure1-1.png",
        "caption": "Fig. 1. Surface parameters of the modified rack cutter with a parabolic profile.",
        "texts": [
            " We also investigate the influence of shaving cutter and work gear pair assembly errors on the topologies, contact ellipses, and transmission errors of the proposed involute helical gear surface. Two numeral examples are presented to illustrate and verify the merits of the proposed gear shaving methodology for longitudinal gear crowning. 2. Mathematical model of a modified shaving cutter Basically, the proposed modified involute helical shaving cutter can be generated by a modified rack cutter with a parabolic tooth profile, as shown in Fig. 1. Hence, the position vector and unit normal vector of the rack cutter's right-hand side profile can be expressed in the coordinate system S8(x8,y8,z8) as follows: r8 \u00bc x8; y8; z8;1\u00bd T \u00bc u1;du 2 1; v1;1 h iT ; \u00f01\u00de and and and n8 \u00bc nx8;ny8;nz8 h iT \u00bc \u22122du1;1;0\u00bd T ; \u00f02\u00de u1 and v1 are the surface parameters, and d is the parabolic coefficient of the modified rack cutter (Fig. 1). where Fig. 2 shows coordinate systems for schematic generation mechanism of the modified helical shaving cutter. The coordinate systems S8(x8,y8,z8), Ss(xs,ys,zs), and S4(x4,y4,z4) in the generation mechanism of the modified involute helical shaving cutter are rigidly connected to themodified rack cutter, involute helical shaving cutter, and frame, respectively. Coordinate systems S7(x7,y7,z7), S6(x6,y6,z6) and S5(x5,y5,z5) are auxiliary coordinate systems for reference. The generated shaving cutter rotates at angle \u03c61 about the z4-axis, while the modified rack cutter translates a distance ro1\u03c61 without rotation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000474_00368791311303456-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000474_00368791311303456-Figure2-1.png",
        "caption": "Figure 2 Geometrical configuration of journal bearing",
        "texts": [
            " In the vicinity of inlet groove, the temperature of recirculating fluid, Trec, is normally higher than the temperature of the incoming supply lubricant, Tsupply. Thus, the recirculating flow transfers a portion of its energy to the supply oil at the inlet. An energy balance at the inlet gives (Khonsari et al., 1996): Tmix \u00bc TrecQrec \u00fe TsupplyQsupply Qrec \u00feQsupply \u00f015\u00de where Qrec and Qsupply are given by the following equations: Qrec \u00bc Z lL 0 Z ri l\u00f0p2ui\u00de Vu\u00f0r;p2 ui ; z\u00dedrdz \u00f016\u00de Qsupply \u00bc Z lL 0 uiviridz \u00f017\u00de where in equation (16) l is the length of line between the center of bush and the journal surface (Figure 2). In addition, at the groove location, the source term in the energy equation is neglected across the lubricant film (Tucker and Keogh, 1995). Furthermore, for all dependent variables, periodic boundary condition in circumferential direction is implemented as follows: \u00f0 p; T; u; v; w; k; 1\u00deu\u00bc0 \u00bc \u00f0 p; T; u; v; w; k; 1\u00deu\u00bc2p \u00f018\u00de \u203a\u00f0 p; T; u; v; w; k; 1\u00de \u203au u\u00bc0 \u00bc \u203a\u00f0 p; T; u; v; w; k; 1\u00de \u203au u\u00bc2p \u00f019\u00de Besides, to calculate the inlet oil velocity vi at each iteration, the lubricant side leakage is computed as: Qside \u00bc Z ri rs Z 2p 0 wrdrdu \u00f020\u00de in which, w is the axial velocity component at the outlet section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000361_iccais.2013.6720561-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000361_iccais.2013.6720561-Figure4-1.png",
        "caption": "Fig. 4. Typical utility-scale wind turbine main components",
        "texts": [
            " There will be a degree of overlapping between the functions of the different subsystems in such a way that some CMS\u2019s subsystems will monitor many parts of the WT. The approach proposed in this review is to differentiate CM techniques applied on WT's subsystem from CM techniques applied on WT's overall system. Subsystem level CM of WTs is based on subcomponents related local parameters [21] [33] [34]. It enables the acquisition of information\u2019s on specific components and therefore, the precise localisation of eventual failures. Typical utility-scale WT main components are presented in Fig.4 while an example of function model for the monitoring of a WECS based on subsystems approach is presented in Fig.5. Subsystems CM can be classified into two main sub categories: those based on destructive tests (DT) and those based on non-destructive tests (NDT). Subsystems CM based on DT uses: vibration analysis, oil analysis, strain measurement, electrical effects, shock pulse method, physical condition of materials, self-diagnosis sensors and others. Subsystems CM based on NDT uses: ultrasonic testing techniques, visual inspection, acoustic emission, thermography, performance Monitoring and radiographic inspection"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000109_1.1713558-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000109_1.1713558-Figure7-1.png",
        "caption": "FIG. 7. Excess metal method. (1) Mold cavity. (2) Ring of excess metal. (3) Mold. (4) ChllJ.",
        "texts": [
            " The formation of columnar crystals in the aniso tropic Alnico alloys has now been known for about 15 years to enhance the magnetic properties, particularly (BH)max, The usual method of producing such crystal growth is to allow the alloy to solidify while heat is withdrawn in one direction only. Radial heat flow may be prevented by casting into a mold which has been preheated in a furnace or heated internally by an exothermic reaction as shown in Fig. 6. Alternative or supplementary methods of preventing radial loss of heat are to fill first with hot metal a ring surrounding the casting as shown in Fig. 7, or merely to arrange ~an:y castings in proximity. Zone melting, illustrated 10 FIg. 8, has also been successfully used; sometimes a susceptor is added to prevent radial heat losses. The use of continuous casting has also been patented. Columnar alloys are still cast only to a small but nevertheless increasing extent. While sometimes the need to use columnar magnets arises from the require ment that the smallest possible magnet should be used for a given purpose, blocks of Columax have also been built up into very powerful magnets"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001168_icuas.2017.7991400-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001168_icuas.2017.7991400-Figure3-1.png",
        "caption": "Fig. 3. Quadcopter scheme with movement axis, thrust forces, motor torques and reference frame.",
        "texts": [
            " The nonlinear model is defined as a function of the state and input vectors X and U as X\u0307 = f(X,U), where the state vector is X = [ x x\u0307 y y\u0307 z z\u0307 \u03c8 \u03c8\u0307 \u03b8 \u03b8\u0307 \u03c6 \u03c6\u0307 ]T , the input vector is set as U = [ u1 u2 u3 u4 ]T =[ u \u03c4\u03c8 \u03c4\u03b8 \u03c4\u03c6 ]T , and u = F1 + F2 + F3 + F4, \u03c4\u03c8 = T2 + T4 \u2212 T1 \u2212 T3, \u03c4\u03b8 = L(F4 + F1 \u2212 F2 \u2212 F3), \u03c4\u03c6 = L(F1 + F2 \u2212 F3 \u2212 F4), (1) where x, y and z represent the quadcopter position in three dimensions, \u03c8 is the yaw angle, \u03b8 is the pitch angle, \u03c6 is the roll angle, u is the throttle input and \u03c4\u03c6, \u03c4\u03b8 and \u03c4\u03c8 the rolling, pitching and yawing moment, Fi is the thrust force applied by the i-th motor, Ti is the torque each motor exert around the z-axis, and L is the distance between the center of mass of the quadcopter and the rotor of the motors. There are used two orthogonal geometrical frames for reference known as body frame (composed by x, y and z) and the inertial frame (composed by xworld, yworld and zworld). These variables are shown in Fig. 3. The equations of motion that define the dynamics of the quadcopter are defined as x\u0308 = u1 m (cos(\u03c6) sin(\u03b8) cos(\u03c8) + sin(\u03c6) sin(\u03c8)) y\u0308 = u1 m (cos(\u03c6) sin(\u03b8) sin(\u03c8)\u2212 sin(\u03c6) cos(\u03c8)) z\u0308 = u1 m (cos(\u03c6) cos(\u03b8))\u2212 g \u03c8\u0308 = \u03c6\u0307\u03b8\u0307 Jxx \u2212 Jyy Jzz + u2 Jzz \u03b8\u0308 = \u03c6\u0307\u03c8\u0307 Jzz \u2212 Jxx Jyy + u3 Jyy \u03c6\u0308 = \u03b8\u0307\u03c8\u0307 Jyy \u2212 Jzz Jxx + u4 Jxx (2) where J\u2022\u2022 are the moments of inertia of the quadcopter around each of its body axes [13], [14]. This is a simplified model of a real quadcopter; a complete quadcopter model can be found in [15]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003482_s00170-018-2324-z-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003482_s00170-018-2324-z-Figure5-1.png",
        "caption": "Fig. 5 Temperature distribution of a workpiece and b tool (rake angle and clearance angle are 15\u00b0 and 20\u00b0)",
        "texts": [
            " This analysis step is used to indicate that a dynamic coupled thermal- stress analysis is to be performed using explicit integration. The cutting simulation of TC21 alloy is shown in Fig. 4. The chip formation and the distribution of stress at different cutting times were obtained through the simulation. As shown in Fig. 4, the primary shear band of chip can be observed clearly. The maximum stress was 1324 MPa and located in the primary shear band. The temperature distribution between the workpiece and tool in the cutting of TC21 alloy is shown in Fig. 5. Figure 5 shows that the chip and tool maximum temperature are 116.6 and 228.7 \u00b0C. And the region of the maximum temperature at tool and chip is in the contact surface between tool and chip and is 0.33 mm from the tip the region. Figures 6 and 7 are the distribution of residual stress in the cut layer of TC21 alloy after oblique cutting. From Fig. 6, it can be seen that the residual stress after cutting has no obvious change with the increasing of rake angle. But in Fig. 7, the residual stress decreased with the increasing of clearance angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000669_1.c031794-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000669_1.c031794-Figure11-1.png",
        "caption": "Fig. 11 The F-18 HARV.",
        "texts": [
            "C 03 17 94 control events due to damage is also being pursued for both civilian and military aircraft [23\u201325]. To assess the potential of the pseudosliding mode approach to contribute to the solution of this problem, it is essential that some form of \u201cpiloted\u201d evaluation of a flight control system incorporating this approach be undertaken. As a preliminary and limited first-step in such an evaluation, a desk-top, human-in-the-loop simulation was conducted using simplified longitudinal and lateral/directional model of the NASA F-18 HARV. The model was taken from [26]. The vehicle, itself, is shown in Fig. 11. The linear model to be studied is given below for a flight condition of Mach No. 0:6, Altitude 30; 000 ft. _x t Ax t Bu t (35) Longitudinal Model: x t t ; q t ; t T A 0:5088 0:994 0 1:131 0:2804 0 0 1:0 0 2 4 3 5 (36) u t HT; PTV T angle of attack, deg q pitch rate, deg =s pitch attitude, deg HT horizontal tail deflection, deg PTV pitch thrust vectoring, deg Fig. 6 Altitude tracking for five uncertainty cases. Fig. 7 Velocity tracking for five uncertainty cases. D ow nl oa de d by K U N G L IG A T E K N IS K A H O G SK O L E N K T H o n Fe br ua ry 1 , 2 01 5 | h ttp :// ar c"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003432_teme-2018-0004-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003432_teme-2018-0004-Figure2-1.png",
        "caption": "Figure 2: Schematic figure of the journal bearing test bench with AE sensor positioned on the bearing back.",
        "texts": [
            " The AE sensor acquires the elastic stress waves and converts them into electrical AE signals. The transducer element is always a piezoelectric crystal which is commonly made of lead zirconate titanate (PZT). The AE signals have to be amplified and filtered and subsequently forwarded to the data acquisition system. All tests were conducted at a journal bearing test bench, which has been specifically developed for this application. These tests refer to investigations on principle and not to tests of a journal bearing located in a turbofan planetary gearbox. Figure 2 shows the journal bearing test bench. A servomotor drives the shaft. The journal bearing consists of a bearing bush, which is made of red bronze (RG7) and a bearing back to prevent dilatation. It is located between two support roller bearings. Two nylon stripes are applied between the shaft and the supporting bearings to reduce interfering signals. A pneumatic cylinder is used to apply load on the bearing back and oil is provided through a drilling. The temperature and oil inlet pressure are held nearly constant at 27 \u00b0C and 2 bar"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001401_j.measurement.2018.01.031-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001401_j.measurement.2018.01.031-Figure2-1.png",
        "caption": "Fig. 2. The power open type gear test machine.",
        "texts": [
            " (3) Substituting the calculation results of friction-related losses at the bearings, total power under loaded conditions and the gear transmission efficiency both unloaded and loaded conditions into the gear meshing efficiency calculation formula, and gear meshing efficiency by measured data from gear test machine is obtained. The flowchart for the calculation of gear meshing efficiency is shown in Fig. 1. Based on these work, we will explain in detail how the experimental gear meshing efficiency is calculated. The gear transmission efficiency test machine can be divided into two categories, i.e., the power open type and power closed type [11]. Figs. 2 and 3 are their typical principle diagrams, respectively. Their transmission efficiencies are respectively calculated as follows. In Fig. 2, gear transmission efficiency is calculated as follows. =\u03b7 T n T n 2 2 1 1 (1) In Fig. 3, the reaction gearbox and the test gearbox are identical. The transmission efficiency of normal-reverse rotation is almost the same. We assume that the power flow direction is clockwise and the supplementary power of the motor is \u0394P. Then: = +P P P\u0394a d (2) =P P \u03b7b a (3) =P P \u03b7d c (4) = \u2217P T n \u03c0\u0394 2 /601 1 (5) = = \u2217P P T n \u03c02 /60c b 2 2 (6) According to Eqs. (2)\u2013(6), we obtain: + \u2212 =T n \u03b7 T n \u03b7 T n 02 2 2 1 1 2 2 (7) Solving the Eq",
            " 2 and 3) \u03b7T gear transmission efficiency under loaded conditions \u03b7S gear transmission efficiency under unloaded conditions PT total power under loaded conditions PS losses in the form of windage and swinging oil PM losses at the gear mesh PB friction-related losses at the bearings \u03b7M gear meshing efficiency \u03bcbi (i= 1, 2) coefficient of friction for the bearing Wbi (i = 1, 2) radial load applied to bearing dborei (i= 1, 2) bore diameter of bearing T torque applied to a gear d diameter of reference circle Ft circumferential force Fr radial force Fn normal force Fa axial force \u03b1n normal pressure angle \u03b2 helix angle = \u2212 + + = \u2212 \u2212 + + = \u2212 \u2212 = \u2212 \u2212 \u2212 \u03b7 P P P P \u03b7 P P P P \u03b7 P P P P \u03b7 \u03b7 P P 1 1 (1 ) (1 ) T S M B T S T M B T S M T B T S M B T (9) According to Eq. (9), we obtain: = + \u2212 +\u03b7 \u03b7 \u03b7 P P 1M T S B T (10) Where, The Eqs. (9) and (10) of the paper are the same as that of Ref. [3], [8], [9] and [10] in the form. But the calculation method of PT, PB, \u03b7T and \u03b7S are completely different. The calculation of \u03b7T and \u03b7S has been described in Section 2. PT and PB are calculated as follows, respectively. In Fig. 2, PT can be represented as, = \u2217P T n \u03c02 60T 1 1 (11) In Fig. 3, PT can be represented as, = \u2217P T n \u03b7 \u03c02 60T T 2 2 (12) PB can be calculated by using the manufacture\u2019s specifications and published bearing efficiency models [12], In Fig. 2, PB can be represented as, = \u2217 + \u2217P \u03bc W d n \u03c0 \u03bc W d n \u03c02 60 2 60B b b bore b b bore1 1 1 1 2 2 2 2 (13) In Fig. 3, PB can be represented as, = \u2217\u239b \u239d \u2217 + \u2217 \u239e \u23a0 P \u03bc W d n \u03c0 \u03bc W d n \u03c02 2 60 2 60B b b bore b b bore1 1 1 1 2 2 2 2 (14) In the paper, the radial load Wb applied to bearing is replaced by the radial force Fr applied to gear. Taking a helical gear as an example, the force analysis of helical gear tooth is shown in Fig. 6. The calculation of radial force Fr is represented in Eqs. (15) and (16) =F T d 2 t (15) =F F \u03b1 \u03b2tan /cosr t n (16) Substituting PT and PB with (10), we can obtain the gear meshing efficiency \u03b7M"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003068_iceee2.2017.7935821-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003068_iceee2.2017.7935821-Figure2-1.png",
        "caption": "Figure 2. Quadcopter hardware",
        "texts": [
            "81 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 B= 0 0 0 0 (4) 0 0 0 0 0 0 0 0 0 . 5917 0 0 0 0 36.7769 0 0 0 0 36. 7769 0 0 0 0 18.9415 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 C=O 0 0 0 0 1 0 0 0 0 0 0 (5) 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 1 In Figure 1, angles and forces of the quadcopter is presented, also in Figure 2, hardware system configuration of the modelled quadcopter is seen. Totally, there are 12 states of the modelled quadcopter system which are (x, y, Z, qJ, 8, 1./1, X, y, i, <jJ, (), l/J). x, y, z are positions along each associated axis. (qJ, 8, 1./1) are angles relative to body frame and following six states are derivatives, respectively. Note that, since there is not any feedforward term in dynamics, D is a zero matrix. There are 4 controllable states in this model while all of the states are observable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001216_ssp.265.535-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001216_ssp.265.535-Figure2-1.png",
        "caption": "Fig. 2. Deposition patterns of part produced by DLD technology \u0430) deposition scheme b) overlay scheme.",
        "texts": [
            " Microstructure and properties of deposited samples were analyzed. Also the influence of hot isostatic pressing (HIP) on mechanical properties are observed. The samples for investigation was produced DLD technology at the Institute of Laser and Welding Technologies SPbPU (ILWT) using laboratory stand based on fiber laser LS-5 with power output 5kW (Fig.1) [20, 21]. Fig.1. Laboratory set-up for study the process of high speed direct laser deposition [20]. Samples are cylindrical rods with radius size 6.5 mm (Fig. 2a). Strategy of patterns deposition spirally from outside to the center with laser beam spot size ds=1.2 mm, moving from point A to point B (movement pattern shown on Fig. 2b). Two passes along the outer diameter, then along a tapering spiral for 3 mm at a speed of 0.7mm/sec, the next 3 mm to the center with a speed of 2 mm/sec. After deposition of the first layer work tool up on height h (layer height), after that according to the scheme Fig.2 deposited next layer. Then that process repeated before obtaining a required height of cylinder. Side nozzle with outlet diameter in the range of 1-2 mm was used for formation of gas powder jet. Was investigate three samples produced on mode with different fractional composition of building material, layer height, gas feed rates and at local protection by argon and at argon environment. Mode parameters is in the table 1. Titanium alloy Ti-6Al-1V-1Mo-2Zr has been used as a material for deposition, powder is shown in Fig. 2 (a,b), the chemical composition is in the Table 2. Fractional (density) composition \u2013 40 \u2013 160 \u00b5m and 160-200, the shape of particles is spherical. In the composition also present particles with irregular shape. Table 2. The chemical composition of alloy Ti-6Al-1V-1Mo-2Zr. Ti Al Mo V Zr Fe Si O other Base 5.5-6.8 0.5-2.0 0.8-1.8 1.5-2.5 0.3 0.15 0.16 0.5 (a) (b) Fig. 3. Powder granule of alloy VT20; (a) fractional composition 40 \u2013 160 \u00b5m (b); fractional composition 160-200 \u00b5m. Metallographic analysis of deposited samples were carried out on microscope DMI 5000 (Leica) with software Tixomet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001956_citcon.2015.7122604-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001956_citcon.2015.7122604-Figure15-1.png",
        "caption": "Fig 15: Insulating Sleeve",
        "texts": [
            " However, one of the simplest methods is to insulate the bearing(s) and disrupt the current flow. This can be done by installing an insulating shield in the bearing housing bore or by insulating the bearing itself. Using the bearing as the insulator is often the easiest and most effective method. Using the bearing as an insulator or disruptor of high frequency current flow can be done in two ways. The first is by adding a ceramic coating to the outer diameter of the outer ring or the bore of the inner ring (Fig. 15). The second is to use rolling elements that are made of ceramic material. Depending on the bearing size, one method may be more economical than the other. The ceramic ring coating is made of aluminum oxide CAlz03) and is nominally 100 microns thick. This coating is a pure insulator in DC motors with a resistance of 50 Mega ohms or more. The dimensions and tolerances of the bearing remain the same as that of a standard bearing without the coating. Figs.I7 & 18. 978-1-4799-5580-0/15/$31.00 \u00a9 IEEE 2014-CIC-1045 In AC motors with high frequencies produced by PWM converters, the ceramic coating or rolling element does not act as a pure insulator or resistor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001592_tie.2018.2868027-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001592_tie.2018.2868027-Figure1-1.png",
        "caption": "Fig. 1. Schematic of a 2-pole TROMAG; (a) 3D view; (b) side view.",
        "texts": [
            " However, the circuits can be readily adapted to other types of magnetic gears and other driven, damped harmonic oscillator energy harvesting applications. Equivalent circuit for the trans-rotary magnetic gear Amin Abolhasani, Student member, IEEE and Siavash Pakdelian, Member, IEEE U 0278-0046 (c) 2018 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. II. DYNAMIC MODEL AND EQUIVALENT CIRCUIT OF THE TROMAG The TROMAG, depicted in Fig. 1, is a magnetic gear tailored to linear motion applications. In the literature, it is also referred to as magnetic screw [6] and magnetic lead screw [8]. It consists of a rotor and a translator which are both made of iron cores covered by helically shaped alternating permanent magnet poles. The translator motion in Z direction results in the rotor rotation about the Z axis, and vice-versa. The gear ratio, G, is defined as the ratio of the translator force, ft, to the rotor torque \u03c4r, and equals G=ft/\u03c4r = 2\u03c0/(P\u03c4P), where \u03c4P is the pole pitch and P is the number of poles as viewed from the XY plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002911_s1068798x16120194-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002911_s1068798x16120194-Figure1-1.png",
        "caption": "Fig. 1. Determination of the forces F(1) and F(2) and cutlayer thickness ak(\u03d5) as a function of the mill\u2019s positional angle \u03d5: (1) mill tooth.",
        "texts": [
            " ters of the speed dependence of the forces; is the modulus of the force at the tool\u2019s rear face with constant longitudinal velocity of the machine tool\u2019s table. This is also a periodic function of the time, with a constant component. Suppose that we specify the steady trajectories of the deformational displacement and the velocity of the mill tooth\u2019s tip in the coordinate system whose motion is characterized by the constant longitudinal velocity of the machine tool\u2019s table and the tool rotation (Fig. 1). On the basis of Eq. (1), they correspond to the trajectory They may be calculated by the rules outlined in [28]. Then the variation in the cut-layer thickness , the cutting speed , and the supply velocity may be written in the form (2) All the functions in Eq. (2), as well as X* and Y*, are specified periodic functions of the time. They consist of constant and periodically varying functions with the same period. In Fig. 1, the bold curve corresponds to variation in kinematic thickness of the cut layer , disregarding the elastic deformational displacement. The dashed curve shows the variation in cut-layer thickness when the elastic displacement for the steady state is taken into account. In Eq. (1), the modulus of the primary force component is (0)( )F t \u2217 \u2217= 1 2 * { , }TX X X v v1 2{ ( ), ( )}t t \u03a9 \u2217 \u2217 \u2217= 1 2{ , } .TY Y Y \u03a3k, ( )a t \u03a3v 2, ( )t \u03a3v1, ( )t \u03a3 \u03a3 \u03a3 \u2217= + = + = + v v v v v v k, k 1 1, sup 1 2, 2 ( ) ( ) ( ); ( ) ( ) ( ); ( ) ( ) ( )"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001146_0954406217722380-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001146_0954406217722380-Figure4-1.png",
        "caption": "Figure 4. Schematic diagram of the (a) rotor system and (b) support B1 and bearing.",
        "texts": [
            " In order to predict the cage motions due to rotor unbalance theoretically, the dynamic models of the rotor system and the bearing cage are presented. Dynamic equations of the rotor system The rotor system for the scaled test rig has two support assemblies including angular contact ball bearing, bearing house, squirrel cage, as shown in Figure 3. The stiffness of support assembly mainly depends on the squirrel cage, because it is much more compliant than other parts. And the stiffness of supports can be obtained by referring to Wang et al.33 The dynamic model for the rotor system is illustrated in Figure 4(a). Its two flexible supports are located at point B1 and B2 on the short rigid shaft and the disc is located at point C. The horizontal and vertical stiffness of point B1 and B2 are k1y, k1z, k2y, k2z. The distances among them are represented as l, a, b, also shown in Figure 4(a). Based on the assumptions that the disc and short shaft are rigid, the two supports are compliant and linear, and the dynamic equations of the rotor system with four-DOFs are deduced with the Lagrangian approach as follows mr \u20acyr \u00fe \u00f0k1y \u00fe k2y\u00de yr \u00fe \u00f0 ak1y \u00fe bk2y\u00de rz \u00bc mere 2 cos\u00f0 t\u00de mr \u20aczr \u00fe \u00f0k1z \u00fe k2z\u00dezr \u00f0 ak1z \u00fe bk2z\u00de ry \u00bc mere 2 sin\u00f0 t\u00de \u00fe Gr Jd \u20ac y \u00fe Jp _ rz \u00fe \u00f0a 2k1y \u00fe b2k2y\u00de ry \u00f0 ak1z \u00fe bk2z\u00dezr \u00bc 0 Jd \u20ac z Jp _ ry \u00fe \u00f0a 2k1z \u00fe b2k2z\u00de rz \u00fe \u00f0 ak1y \u00fe bk2y\u00de yr \u00bc 0 \u00f01\u00de where mr and Gr denote the rotor mass and gravity",
            " mere 2 is the amplitude of the time-varying force due to rotor unbalance. By solving the above dynamic equations numerically, the unbalance responses of rotor can be estimated. Remembering the rigid shaft assumption, the displacements of the support B1 can be easily computed. So the time-varying reaction forces of support B1 (Fy1, Fz1) and B2 (Fy2, Fz2) due to unbalance are obtained as follows Fy1 \u00bc k1y\u00f0 yr a rz\u00de Fz1 \u00bc k1z\u00f0zr \u00fe a ry\u00de Fy2 \u00bc k2y\u00f0 yr \u00fe b rz\u00de Fz2 \u00bc k2z\u00f0zr b ry\u00de \u00f02\u00de The ball bearing and support B1 are assumed to be in static equilibrium illustrated in Figure 4(b), so the forces of bearing in the support B1 are calculated as follows Fy \u00bc Fy1 Fz \u00bc Fz1 \u00f03\u00de Dynamic model of bearing cage A dynamic model of the bearing cage with threeDOFs is developed, in which the interactions among the balls, inner rings and cage are taken into account, and the reaction forces due to rotor unbalance are also introduced as boundary conditions of bearing. 1. Dynamic equations of cage The cage motions are represented by its mass center oc in the inertial frame oyz described by the vector (yc, zc) and its rotational angle c"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002235_cgncc.2014.7007533-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002235_cgncc.2014.7007533-Figure2-1.png",
        "caption": "Fig. 2. Inertia and body coordinate system of the quadrotor",
        "texts": [
            " The quadrotor helicopter used in this paper is developed based on DJI-450 framework as shown in Fig. 1. The quadrotor UAV consisting of four fixed pitch angle rotors powered by four electronic motors and mounted at the end of an X-shaped frame. The micro UAV system is composed four 978-1-4799-4699-0/14/$31.00\u00a92014 IEEE subsystems, the airframe, the power system and the flight control system. To improve the safety of experiments, carbonfiber tubes are mounted on the airframe[7]. The derivation of the dynamics is based on inertial and body frames. And the two frames are defined as shown in Fig. 2. Let ( , , )B x y z be the frame attached to the vehicle and ( , , )E x y z denote the inertial frame. The origin of the inertial frame is set to the take-off point while the origin of body frame is in the vehicle center of gravity (CG). The rotation matrix ( , , )R \u03c6 \u03b8 \u03d5 from body frame to inertial frame depends on the three Euler angles ( , , )\u03c6 \u03b8 \u03d5 representing, respectively, the roll, the pitch and the yaw. To simplify the model process and facilitate the flight control design, the quadrotor is seen as a rigid-body and the center of gravity coincides with the origin of the body coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003688_042044-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003688_042044-Figure5-1.png",
        "caption": "Figure 5. Schematics of the specimens cut off from articles 1 and 2 \u2013 specimens cut off along the building direction; 3 and 4 \u2013 specimens cut off transversal to building direction, 5 \u2013 article, 6 \u2013 deposition trajectory, 7 \u2013 substrate.",
        "texts": [
            " The article\u2019s surface is rather smooth without any unmelt wire segments or excess melting the edges (figure 2c). This article may be acceptable from the viewpoint of the layers deposited but has curved walls, probably resulted from their insufficient thickness. Article #3 was obtained at 0.18 m/s (figure 2d, figure 3) and characterized by smooth surfaces with- out any unmelt wire or excess melting. 4 1234567890 \u2018\u2019\u201c\u201d The mechanical characterization of the articles obtained was performed using compression tests on specimens cut from the articles as shown in figure 5. The Testsystems 110M-10 test machine has been used in the process at ambient temperatures. Table 2 contains mean yield stress values for different specimens tested. It follows from the results that article #3 shows the maximum stable mechanical strength distribution. Article #2 is characterized by minimum yield stress in the direction of deposition trajectory. 1234567890 \u2018\u2019\u201c\u201d 6 1234567890 \u2018\u2019\u201c\u201d The article #1 (top part) is characterized by the yield strength values higher in the building direction as compared to that of transversal direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003128_miltechs.2017.7988729-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003128_miltechs.2017.7988729-Figure7-1.png",
        "caption": "Figure 7. Coordinate systems at center of mass of bodies 4, 5 and 6",
        "texts": [
            " The shooter \u2013 weapon system includes nine rigid bodies: \u2022 body 1 \u2013 shooter shoulder of mass m1, \u2022 body 2 \u2013 weapon frame of mass m2, \u2022 body 3 \u2013 breech system of mass m3, \u2022 body 4 \u2013 upper right arm (APBP) of mass m4, \u2022 body 5 \u2013 lower right arm (BPCP) of mass m5, \u2022 body 6 \u2013 right hand (CPDP) of mass m6, \u2022 body 7 \u2013 upper left arm (ATBT) of mass m7, \u2022 body 8 \u2013 lower left arm (BTCT) of mass m8, \u2022 body 9 \u2013 left hand (CTDT ) of mass m9 and It is necessary to take into consideration the mass mi of weapon moving parts. In order to analyze the dynamics of the shooter \u2013 weapon system, Descartes coordinate systems have been established at the mass center of each bodies and the whole system as shown in Fig. 6 and Fig. 7, where: O0 represents the stationary coordinate system fixed on the ground, Ok accounts for the local coordinate system established at the center of mass of k-th rigid body, { }for = 1, 2,..., 9k . The coordinates of system rigid bodies are labeled as follows: 1q - longitudinal displacement of body 1 along X0-axis, 2q - angular displacement of body 2 about X1-axis, 3q - angular displacement of body 2 about Y1-axis, 4q - angular displacement of body 2 about Z1-axis, 5q - longitudinal displacement of body 3 along X2-axis, 6q - angular displacement of body 4 about X1-axis, 7q - angular displacement of body 4 about Y1-axis, 8q - angular displacement of body 4 about Z1-axis, 9q - angular displacement of body 5 about Z4-axis, 10q - angular displacement of body 7 about X1-axis, 11q - angular displacement of body 7 about Y1-axis, 12q - angular displacement of body 7 about Z1-axis, 13q - angular displacement of body 8 about Z7-axis, 14q - angular displacement of body 6 about X5-axis, 15q - angular displacement of body 6 about Y5-axis, 16q - angular displacement of body 6 about Z5-axis, 17q - longitudinal displacement of body 6 along X2-axis, 18q - angular displacement of body 9 about X8-axis, 19q - angular displacement of body 9 about Y8-axis, 20q - angular displacement of body 9 about Z8-axis, 21q - longitudinal displacement of body 9 along X2-axis, iq - longitudinal displacement of body i along Xi-axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000307_j.jallcom.2013.05.119-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000307_j.jallcom.2013.05.119-Figure5-1.png",
        "caption": "Fig. 5. Microhardness distribution of Ni\u201333 wt.%Sn hypereutectic alloy after laser remelting (V = 0.1 mm/s).",
        "texts": [
            " 4a and b shows the patterns of Ni3Sn compound and a-Ni solid solution, which are indexed separately. Fig. 4c and d shows the pole figures of Ni3Sn and Ni phase respectively. In the EBSD map indexed to Ni3Sn phase only, this region can be divided into three subregions, i.e., two anomalous eutectic regions and one lamellar eutectic region (Fig. 4a). The Ni3Sn phase shows one color in each of subregions. In the EBSD map indexed to a-Ni phase only, the Ni phase shows a colorful pattern, except in the ellipse region and rectangle region marked in Fig. 5b. The pole figures of Ni3Sn phase show that Ni3Sn grains are oriented along three major poles of the {0001}-type (Fig. 4c), but the pole figures of a-Ni phase show that a-Ni grains are oriented along many major poles of the {111}-type (Fig. 4d). The result of EBSD maps and the pole figures shows that Ni3Sn phase is continuous and well oriented in the individual colonies, whereas the a-Ni phase is discontinuous and inconsistent in orientation. This results are in agreement with the previous studies by using high undercooling method [9,18]",
            " The Fig. 4. EBSD analysis of the anomalous eutectic structure at the bottom of the specimen. pole figure of Ni3Sn; and (d) {111} pole figure of Ni phase. presence of coarsening irregular coupled eutectic growth around the anomalous eutectic means that both coupled growth and decoupled growth can result in the anomalous eutectic formation to a certain extent. The microhardness distributions of Ni\u201333 wt.%Sn hypereutectic alloy after laser remelting with the scanning speed of 0.1 mm/s are shown in Fig. 5. Microhardness evolution was measured along vertical lines starting from the top (lamellar eutectic plus primary Ni3Sn zone) of the sample and moving across the central part (lamellar eutectic zone) and the bottom part (anomalous eutectic zone). Obviously, the microhardness decreased from the top, the central and finally to the bottom. This can be attributed to the coarse anomalous eutectic microstructure at the bottom and finer regular (a) EBSD pattern indexed to Ni3Sn; (b) EBSD pattern indexed to Ni phase; (c) {0001} lamellar eutectic microstructure at the central"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002772_j.ifacol.2016.07.125-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002772_j.ifacol.2016.07.125-Figure1-1.png",
        "caption": "Fig. 1. VTOL configuration.",
        "texts": [
            " Notations: The notations used in this article are standard. For a matrix A, A and A\u22121 denote its transpose and inverse respectively. He{A} denotes A + A . In the following, the explicit time dependence of the variables will be omitted. 3. PVTOL DYNAMICS The PVTOL is composed of two independent motors with two propellers that produce a force fi and torque \u03c4\u03b8 in the direction of the propellers axis of rotation. The propellers 1 turns in counterclockwise direction, while the propellers 2 turns in clockwise direction as shown in Fig. 1. Vertical displacement on the z axis is generated increasing or decreasing, by the same magnitude, the velocity in the motors. Moreover, inversely changing the speed of the motors 1 and 2, a roll rotation \u03b8 coupled with a lateral motion in the x axis is produced. E = {Oe, xe, ze} denotes an inertial frame, and Q = {Oq, xq, zq} denotes a rigid frame attached to center of mass of the vehicle. in this paper, the following is considered: \u2022 The VTOL structure is rigid and symmetrical, \u2022 The center of mass and Oq coincides, \u2022 The propellers are rigid, \u2022 Thrust and drag are proportional to the square of the propellers speed, \u2022 The roll moment create a force that is perpendicular to z axis, \u2022 The yaw dynamics is ignored"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000444_0954406214524743-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000444_0954406214524743-Figure1-1.png",
        "caption": "Figure 1. Piston-connecting rod system. (a) Oversized bigend bearing clearance; (b) hydrodynamic forces at the big-end bearing.",
        "texts": [
            " As found from the experiments, the raw vibration signals on the surface of the engine block are typical second order cyclostationary signals and an appropriate signal processing approach for the diagnosis of bearing knock faults is thus envelope analysis of the vibration signals. Therefore, in order to evaluate the simulation model, the squared envelopes of the simulated acceleration were calculated and compared with the experimental squared envelopes. Principle for the simulation of bearing knock faults If an oversized clearance is considered in the big-end bearing, the piston and connecting rod can be modelled as one motion subsystem and the crank can be modelled as another motion subsystem, as shown in Figure 1. The length of connecting rod is b. The length from the centre of the small end to the centre of gravity (CG) of the connecting rod is b1 and the length from the centre of big end to the CG of the connecting rod is b2. The radius of the crank is R. The coordinate of the crank journal centre is (yc, zc) and the coordinate of the bearing centre is (yb, zb). The friction force on the piston has been considered negligible compared to the combustion force and the vertical inertia force of the piston"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003650_s12206-018-0907-0-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003650_s12206-018-0907-0-Figure2-1.png",
        "caption": "Fig. 2. Displacements and rotations in cylindrical coordinate system.",
        "texts": [
            "n n n nb c w b c\u00e9 \u00f9 \u00e9 \u00f9 \u00e9 \u00f9 \u00e9 \u00f9+ = +\u00eb \u00fb \u00eb \u00fb \u00eb \u00fb \u00eb \u00fbA A B B (10) The vectors in Eq. (10) are defined as { } { } { } { } 1 2 1 2 1 2 1 2 1 2 1 2 1 2 1 2 1 2 1 2 . Tco co co co co co co co co co co co co co co P P P P P Tco co co co co co co co co co P P P P P P P P P P a a a b b b c c c d d d e e e a a b b c c d d e e b c + + + + + + + + + + = \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 \u00d7 = (11) Matrices [A1(n)], [A2(n)], [B1(n)] and [B2(n)] correspond to the circumferential wave number n. 2.2 Cylindrical shell segment The equations of motion of cylindrical shell as shown in Fig. 2 are 21x xN N hU x R q w r q \u00b6 \u00b6 + = - \u00b6 \u00b6 (12a) 21xN N Q hV x R R q q q w r q \u00b6 \u00b6 + + = - \u00b6 \u00b6 (12b) 21xQ Q N hW x R R q q w r q \u00b6 \u00b6 + - = - \u00b6 \u00b6 (12c) 3 21 12 x x x M M hQ x R q rw F q \u00b6 \u00b6 + - = - \u00b6 \u00b6 (12d) 3 21 12 xM M hQ x R q q q rw Y q \u00b6 \u00b6 + - = - \u00b6 \u00b6 (12e) assuming simple harmonic motions with the effects of transverse shear and rotary inertia taken into account. U, V and W are the displacements in the axial, circumferential and normal directions, respectively. F and Y are the rotations in the axial and circumferential directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000639_acc.2015.7171957-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000639_acc.2015.7171957-Figure1-1.png",
        "caption": "Fig. 1. Simplified two dimensional Segway model",
        "texts": [],
        "surrounding_texts": [
            "In this section we will use the OVP controller and a LQR controller to stabilize the system around its unstable equilibrium point z = [0, 0, 0, 0] and compare both controllers regarding performance and robustness. For the design of the OVP controller, we notice that the system (54) satisfies all assumptions of Theorem (Thm. 4.1). Computing \u2207f0 := \u2207 f |x=0 and g(0), \u2207f0 =  0 1 0 0 13.5687 0 0 0 0 0 0 1 \u22125.5678 0 0 0  ,g(0) =  0 \u22120.063 0 0.0546  (55) we see that the pair [\u2207f0, g(0)] is controllable. Using our Theorem (Thm. 4.1), we know that the corresponding pair [A,b] is controllable and thus we can make the system minimum-phase with ys = \u03b8Tz, by making A \u2212 bKT Hurwitz and parametrizing \u03b8 as (37). To design K we apply the LQR approach described in section (IV-C) with \u03bb = 10 and the weighting matrices Qe1 =  100 0 0 0 0 1 0 0 0 0 100 0 0 0 0 1  ,Re1 = 0.1. Parametrizing \u03b8 with the resulting K as (37) leads to the following output vector \u03b8: \u03b8T = [ \u221264.29 \u221225.67 \u22123.16 \u221211.30 ] (56) As a basis for comparison, we will choose the same matrices Qe1 and Re1 for designing the LQR-controller. Thereby, we can assume for the comparison that both controllers are trying to achieve the same performance goal. Using Q = Qe1 and R = Re1 the standard LQR-feedback controller uLQR = \u2212KLQRz is parametrized with the gain matrix KT LQR = [ \u2212636.78 \u2212196.09 \u221231.63 \u221265.70 ] . By simulating the system in closed-loop with both controllers, we obtain the results demonstrated in Fig. 2. Comparing both controller performances with the same initial conditions, we notice that the OVP and LQR controller stabilize the system in a similar manner. On the other hand, when estimating the biggest region of attraction (RoA) for both controllers with respect to \u03c8, we notice that the OVP controller can stabilize the system with larger initial deviations from the equilibrium (up to \u03c8(0) = 90\u25e6), while also providing a better damped closed-loop performance than the LQR-controller. As shown in Fig. 2 the LQR manages to stabilize the system only up to \u03c8(0) = 81\u25e6 and makes the system unstable for any bigger initial \u03c8-deviations."
        ]
    },
    {
        "image_filename": "designv11_13_0003650_s12206-018-0907-0-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003650_s12206-018-0907-0-Figure1-1.png",
        "caption": "Fig. 1. Displacements and rotations in the conical coordinate system.",
        "texts": [
            "1 Conical shell segment The equations of motion of conical shell with the effects of transverse shear and rotary inertia taken into account are 21 sin x x xN N N N hU x x x q q w r a q \u00b6 \u00b6 - + + = - \u00b6 \u00b6 (1a) 21 2 sin tan x xN N N Q hV x x x x q q q q w r a q a \u00b6 \u00b6 + + + = - \u00b6 \u00b6 (1b) 21 sin tan x xQ Q Q N hW x x x x q q w r a q a \u00b6 \u00b6 + + - = - \u00b6 \u00b6 (1c) 3 21 sin 12 x x x x M M M M hQ x x x q q rw F a q \u00b6 \u00b6 - + + - = - \u00b6 \u00b6 (1d) 3 21 2 sin 12 x xM M M hQ x x x q q q q rw Y a q \u00b6 \u00b6 + + - = - \u00b6 \u00b6 (1e) for simple harmonic motions at natural frequency w in radian/second. U, V and W are the displacements in the meridian, circumferential and normal directions, respectively, as shown in Fig. 1. F and Y are the rotations with respect to the circumferential and meridian directions. a is the semi apex angle and x is the distance from the apex O. h and r are the wall thickness and the density of the shell. The stress resultants Nx, N , Nx , Qx, Q , Mx, M and Mx are defined as 1 sin tanx U V U WN C x x x x n a q a \u00ec \u00fc\u00b6 \u00b6\u00e6 \u00f6= + + +\u00ed \u00fd\u00e7 \u00f7\u00b6 \u00b6\u00e8 \u00f8\u00ee \u00fe (2a) 1 sin tan V U W UN C x x x xq n a q a \u00b6 \u00b6\u00e6 \u00f6= + + +\u00e7 \u00f7\u00b6 \u00b6\u00e8 \u00f8 (2b) 1 1 2 sinx x V V UN N C x x xq q n a q - \u00b6 \u00b6\u00e6 \u00f6= = - +\u00e7 \u00f7\u00b6 \u00b6\u00e8 \u00f8 (2c) 1 sinxM D x x x F Y Fn a q \u00ec \u00fc\u00b6 \u00b6\u00e6 \u00f6= + +\u00ed \u00fd\u00e7 \u00f7\u00b6 \u00b6\u00e8 \u00f8\u00ee \u00fe (2d) 1 sin M D x x xq Y F Fn a q \u00b6 \u00b6\u00e6 \u00f6= + +\u00e7 \u00f7\u00b6 \u00b6\u00e8 \u00f8 (2e) 1 1 2 sinx xM M D x x xq q n Y Y F a q - \u00b6 \u00b6\u00e6 \u00f6= = - +\u00e7 \u00f7\u00b6 \u00b6\u00e8 \u00f8 (2f) x WQ Gh x k F \u00b6\u00e6 \u00f6= +\u00e7 \u00f7\u00b6\u00e8 \u00f8 (2g) 1 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003887_j.ymssp.2019.04.039-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003887_j.ymssp.2019.04.039-Figure1-1.png",
        "caption": "Fig. 1. Excitations, transfer paths of planetary gear system (fixed ring gear).",
        "texts": [
            " The modeling of the vibration signal targets the planetary-gear system comprised of one ring gear (r), one sun gear (s), one rotating carrier (c), and N planet gears (pi, i = 1,2..N). The vibration signal refers to the vertical response of the ring gear at the angular position of u \u00bc p=2 where the stationary transducer is mounted. The dominant excitations of the responses of a planetary gear system refer to the gear meshing forces. There are three major transfer paths transferring the excitations to the responses (cf. Fig. 1). Path 1 for planet-ring gear meshing forces: planet-ring meshing location? ring gear? transducer; Path 2 for planet-sun gear meshing forces: planet-sun meshing location? planet gear body? ring gear? transducer; Path 3 for planet-sum gear meshing forces: planet-sun gear meshing location? sun gear? sun gear shaft? sun gear shaft bearing? rear cover? ring gear? transducer [12]. Path 1 is specific for the excitation from the planet-ringmeshing forces. Path 2 is specific for excitation from the planet-sun meshing forces",
            " For the proposed approach, separate consideration of Path 2 is not required due to the following reasons. The gear meshing forces serving as the excitations are obtained with the gear dynamics based lumped parameter model. The lumped parameter model considers the degrees of freedom of the complete gear system including the ring gear, the planet gears, the sun gear and their dynamics coupling effects. In this context, the planet-ringmeshing forces come from the interaction between the ring gear and the subsystem comprised of the carrier, the planet gears, and the sun gear. (cf. Fig. 1), and the planet\u2013sun meshing forces become the internal forces of the sub-system, and these internal forces cannot bypass planet-ring meshing to influence the response signal by Path 2.Meanwhile, governingby the coupling effect the response signal reflects the influence of the planet-sunmeshing forces through the planet-ring gear meshing forces. Moreover, Path 3 is not considered in this study because the path is composed of a chain of gear parts with strong damping effects. In general, Path 1 becomes the focus of this study",
            " This section firstly derives the continuous space-time models for the individual response signal excited by the individual planet-ring gear meshing force, and for the gross response signal excited by all the planet-ring gear meshing forces; secondly interprets the spectral features and the mechanism of the response signal; lastly discusses the relationship of the suggested model versus the existing models. The individual response refers to the response excited by the individual excitation f vi . The excitation refers to the vertical component decomposed from the ith planet-ring gear meshing force f i (cf. Fig. 1), which is represented by f vi kT\u00f0 \u00de \u00bc f i kT\u00f0 \u00decos ui kT\u00f0 \u00de - a\u00bd ; k \u00bc 0;1;2:::\u00f0 \u00de \u00f01\u00de where T denotes a tiny time step; cos ui kT\u00f0 \u00de - a\u00bd represents the projection function to decompose the vertical component from the meshing force. a denotes the meshing pressure angle, ui denotes the angular displacement of the ith planet gear. Let h ui kT\u00f0 \u00de; t\u00bd denote the impulse response function (IRF) of the transfer path for the excitation from the ith planet-ring meshing force. The two-dimensional IRF contains spatial dimension u and time dimension t"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000636_j.proeng.2012.04.032-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000636_j.proeng.2012.04.032-Figure1-1.png",
        "caption": "Fig. 1. Geometry for the scattering experiments, with the arrows indicating the positive direction for the various quantities",
        "texts": [
            " Low-speed ball-bat collisions have shown that the often-made sliding-to-rolling assumption is not valid [1]. Instead, the ball grips the surface of the bat, stretching the ball in the transverse direction, resulting in a significantly higher spin than can be obtained by rolling. The object of the present experiment is to extend the previous studies to much higher speed. * Corresponding author. Tel.: +1-217-333-0965; fax: +1-217-333-1215. E-mail address: a-nathan@illinois.edu. Two different experiments were performed, both with the geometry shown in Fig. 1. A baseball is projected horizontally onto a three-inch diameter wood cylinder that is rigidly bolted to a rigid surface so that no recoil is possible. The incident velocity vector v1 is offset from the centerline of the cylinder by an amount E so that it makes an angle with the normal to the surface of the cylinder. The ball scatters through an angle and has final velocity v2. The incoming and outgoing spins are 1 and 2, respectively, with the positive direction along the outward normal to Fig. 1, corresponding to incident topspin and outgoing backspin. In the first experiment, the ball was projected with a two-wheel pitching machine for which the incident speed and spin were independently adjusted to be in the range 85-110 mph and 1000-3000 rpm, respectively. In the second experiment, the ball was projected without spin at a fixed speed of 120 mph from a pressurized air cannon. In both experiments, the scattering was viewed by a video camera operating at 2000 frames/sec. Care was taken to assure the camera axis was normal to the scattering plane",
            " It was confirmed that the spin axis coincided with the normal to the scattering plane to better than 3 degrees. Neither nor E are directly measured. A similar but slightly different procedure was used for the other experiment. A total of 52 impacts were analyzed. The goal of the analysis is to interpret the experimental results in the context of the collision model described in detail by Cross [3]. First we denote x and y as directions transverse and normal to the surface of the bat at the contact point, respectively (see Fig. 1). The model is described by three parameters. One of these is the coefficient of restitution (COR) for the normal velocity component, 11 2 1 2 1 1 1 10.4 1 1.4 1.4 1.4 11 0.4 1 = + 0.4 1.4 1.4 1.4 y yx x x yx x x x y e ve e v D r r r ee e D v v r v r 2 2 1 1 x x x v r e v r 2 2 1 1 1(1 ) ,x x y yI mv r I mv r mD e v ey=-vy2/vy1, which is approximately 0.5 for a baseball on a rigid wood surface. Another is the tangential COR ex, defined as the negative ratio of final to initial tangential surface velocities of the ball: (1) where r is the ball radius",
            " For larger values, gross slip occurs and the impulse ratio saturates at the value 0.15. The self-consistency among these plots gives us confidence in our interpretation of the data. We next investigate the implications of our results for batted baseballs, requiring modifications to the formalism for a free (rather than clamped) bat that is swung (rather than at rest), as developed in [1]. In the calculation, a baseball is projected horizontally at 85 mph onto a bat that is swung horizontally at 70 mph. The impact parameter E (see Fig. 1) is varied and the post-impact velocity and spin are calculated using our formalism and the fitted parameters. Fig. 3 shows 2 vs. the launch angle for 1= 2000 rpm and for ex=0.3 (solid) and 0 (dashed), respectively. For ex=0.3, the two solid curves are essentially indistinguishable for impacts for which there is no gross slip, showing that the batted ball spin is not strongly dependent on the pitched ball spin for a given launch angle . This conclusion is very different from that found with ex=0 [4,5], as seen from the dashed curves"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000317_s12541-012-0115-2-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000317_s12541-012-0115-2-Figure2-1.png",
        "caption": "Fig. 2 The model of a motion stage driven by a servo motor",
        "texts": [
            " The disturbance observers are designed to decrease the mismatched dynamics between a nominal model and an actual plant with nonlinear properties and to reject instantly disturbances for each stage. In addition, the synchronous controller is designed from the view point of accurate synchronization and reduction of the maximum value of the synchronous error through improvement of damping characteristics for the synchronous control system. Finally, experiment results show that the proposed control system is effective for the skew disturbances. A motion stage is a transport apparatus which converts a rotational motion of a servo motor into a linear motion of a table. Fig. 2 illustrates the model of a motion stage in which the table is driven by a servo motor through a ball screw along the linear guide. The rotational motion and electric circuit of the servo motor which rotates the ball screw are expressed as J \u03b8 t B \u03b8 t T t T t K i t ,\u00a0 (1) u t L \u0131 t R i t K \u03b8 t /K ,\u00a0 (2) where T t is the motor torque, T t is the load torque, \u03b8 t is the angular displacement, i t is the armature current, u t is the voltage input, J is the rotational inertia, B is the viscous friction coefficient, K is the motor driver gain, K is the torque constant, K is the back-emf constant, R is the resistance, and L is the inductance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003324_j.jmapro.2018.01.005-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003324_j.jmapro.2018.01.005-Figure6-1.png",
        "caption": "Fig. 6. Distribution of effective stress in A-A section.",
        "texts": [
            " The stress in bridge nd boss regions was identified as compressive stress. The distribuion of effective stress in forging die was decomposed in different irections, as shown in Fig. 5. Bridge regions were mainly suffered rom compressive stress in Z direction and shear stress in X and Y irections, boss regions were suffered from compressive stress in , Y, and Z directions, groove regions and indent of bridge regions uffered from tensile stress. Distribution of effective stress in A-A ection is shown in Fig. 6. Stress decreased with increasing depth. s the depth of high-stress regions was limited, the performance of orging die could be improved by changing the materials of surface ayers in these regions. The changing values of primary thermal tress during 5 passes forging at the 6 points selected in Fig. 2b re shown in Fig. 7. The tendencies of stress are similar with these of temperatures shown in Fig. 3b. The values of primary thermal stress decreased with increasing depth. The primary thermal stress at P1 excessed 300 MPa and the fluctuation stress range excessed 100 MPa"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001427_978-3-319-89480-5_13-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001427_978-3-319-89480-5_13-Figure2-1.png",
        "caption": "Fig. 2 A rectangular EBM 718 specimen, sample coordinate system, cutting paths, and final cut/ ground/polished specimens for the oxidation studies",
        "texts": [
            " T ab le 1 C he m ic al co m po si tio n of th e po w de r an d as -r ec ei ve d w ro ug ht 71 8 (a ll in w t% ) N i C o C r M o T i M n N b B P T a A l Fe Si S C Po w de r 54 .1 0. 04 19 .0 3. 0 1. 0 0. 1 4. 9 0. 00 1 0. 00 4 < 0. 01 0. 5 B al . 0. 1 < 0. 00 1 0. 03 A sre ce iv ed w ro ug ht 52 .4 0. 05 19 .1 3. 1 0. 8 0. 2 5. 2 0. 00 1 0. 00 3 < 0. 01 0. 6 B al . 0. 2 < 0. 00 1 0. 04 An Arcam A2X EBM machine (Arcam, M\u00f6lndal, Sweden) with an acceleration voltage of 60 kV was used to produce rectangular specimens (\u223c20 mm \u00d7 50 mm \u00d7 3 mm thick), see Fig. 2. Process parameters proposed by Arcam for Alloy 718 was used together with an operating temperature of \u223c1023 \u00b0C and a layer thickness of 75 \u03bcm. A vacuum of 0.2 Pa was used during operation together with a small amount of He in the build chamber. The EBM process parameters used are given in Table 2. The specimens were cut to size (\u223c10 mm \u00d7 10 mm \u00d7 3 mm) and all the surfaces were wet ground using SiC paper up to 2500 grit and polished with a 0.02 \u00b5m silica suspension to produce a smooth surface (Ra < 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000500_s11665-012-0227-y-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000500_s11665-012-0227-y-Figure11-1.png",
        "caption": "Fig. 11 Choosing the representative sections of bent-tube",
        "texts": [
            " Therefore, the springback prediction precision depends greatly on the stress simulated in the bending process and the retracting mandrel process, which has to be studied thoroughly. The distribution and variation of the stress were obtained using the FE model established above. The distribution of the effective stress of the thin-walled rectangular 3A21 tube is shown in Fig. 10. It can be seen that retracting mandrel and springback has released a large amount of the effective stress. Some representative sections are chosen to analyze the tangential stress more clearly. Figure 11 shows the represen- tative nodes A ~ J and A\u2032 ~ J\u2032 on the edges of intrados and extrados in the bending part when the tube is bent to 90\u00b0. J-J\u2032 is the initial bending section, and A-A\u2032 is defined as a reference section which remains unchanged in the bending process. B-B\u2032 ~ I-I\u2032 are 10\u00b0~ 80\u00b0 to the reference section A-A\u2032, respectively. Figure 12 shows the tangential stress of representative nodes A ~ J and A\u2032 ~ J\u2032 on the edges of intrados and extrados after bending and retracting mandrel. It can be seen that there is more tangential compressive stress in the inner side of bent-tube after bending"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002429_1.4031994-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002429_1.4031994-Figure2-1.png",
        "caption": "Fig. 2 Schematic of a sinusoidal asperity loaded with a rigid flat",
        "texts": [
            " The contact radius is related to the noncontact width by a \u00bc k 2 b (15) where a is the contact radius, k/2 is the radius of the sinusoidal asperity, and b is the length of the noncontact region. Then by substitution into Westergaard\u2019s model, the following approximate equation is found: A2 \u00bc p k 2 k p 1 p p 1=2 \" #2 (16) 2.4 FE Model. Figure 1 shows the 3D view of the surface of the sinusoidal asperity. For clarification of the geometry, the amplitude of the sinusoidal asperity shown is D\u00bc 0.5 mm instead of 0.01 mm as modeled. Figure 2 presents a cross section of an axisymmetric sinusoidal surface loaded against a rigid flat. Due to the symmetry along the axis of revolution, only the right-half of the cross section within the period r\u00bc [ R1, R1] is used in the FE model. PLANE183, an eight-node plane element with midnodes, is used for meshing of the entire sinusoidal surface and its substrate. By performing a mesh convergence though gradually increasing the nodal density (especially, the one closest to the contact interface), it is observed from Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000017_icsmc.2012.6378165-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000017_icsmc.2012.6378165-Figure7-1.png",
        "caption": "Fig. 7. Balancing strategies",
        "texts": [
            " To explain the basic idea clearly, the strategies are considered under the following constraints. \u2022 The ground, where the foot/feet of Roboray is on, is flat and not inclined. \u2022 Both yaw joints in the hip are fixed in forward direction. \u2022 The orientation of trunk is controlled to keep the upper body upright based on the orientation sensed by IMU. The control torque for each leg is calculated separately. We use 6 \u00d7 1 vectors such as \u03b8L = [\u03b8L1, \u03b8L2, ..., \u03b8L6] , \u03c4L = [\u03c4L1, \u03c4L2, ..., \u03c4L6] for the 6 joints of left leg. The components of a vector is numbered following the numbers as in Fig. 7(a). The same notations are used for right leg with a subscript R. \u2022 Balancing with a single leg : We consider the case of a robot balancing with a left leg as in Fig. 7(a). Balancing with a single leg is nearly the same as balancing control of the 1-mass pendulum. It is to control the inverted pendulum with the same mass of the robot\u2019s total mass (mtotal), which is located at CoM as shown in Fig. 7(a). The necessary torque is computed as the following. First, the position of target CoM is placed at the ankle joint of a supporting leg in the xyplane and the height of target CoM is set to be h. Then the target position of CoM is (0, 0, h) with respect to the ankle. If we choose the gravity scale factor k as 1 and the virtual gravity vector G\u2032 as [0, 0,\u2212g] , then the control torque of virtual gravity control becomes \u03c4g,L = 2g(\u03b8L). Note that the virtual gravity vector is the opposite to the gravity vector",
            ".., dL6)\u03b8\u0307L and DR(\u03b8\u0307R) = diag(dR1, dR2, ..., dR6)\u03b8\u0307R. The damping coefficients (dL1 \u223c dL6, dR1 \u223c dR6) are determined by experiment. Finally, the desired torque for each leg to balance with left leg is \u03c4d,L = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u03c4g,L1 \u2212 dL1\u03b8\u0307L1 \u03c4g,L2 \u2212 dL2\u03b8\u0307L2 \u03c4g,L3 \u2212 dL3\u03b8\u0307L3 \u03c4pitch \u2212 dL4\u03b8\u0307L4 \u03c4roll \u2212 dL5\u03b8\u0307L5 \u03c4yaw \u2212 dL6\u03b8\u0307L6 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 , \u03c4d,R = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \u2212dR1\u03b8\u0307R1 \u2212dR2\u03b8\u0307R2 \u2212dR3\u03b8\u0307R3 \u2212dR4\u03b8\u0307R4 \u2212dR5\u03b8\u0307R5 \u2212dR6\u03b8\u0307R6 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 . (16) \u2022 Balancing with both legs : The balancing control with two legs is depicted in Fig. 7(b). The procedure to calculate the desired torques for both legs is similar to the above procedure for a single stance leg. Once the location of each foot is known, we set the desired position of CoM at the middle of two feet and the height at h. Let the desired position of CoM from the right feet be (xR, yR, h), then the desired position of CoM from the left leg is (\u2212xR,\u2212yR, h). Then, for the scale factor k = 1, the virtual gravity vector for each leg G\u2032 R = \u2212g\u221a xR 2 + yR 2 + h2 [xR, yR, h] , (17) G\u2032 L = \u2212g\u221a xR 2 + yR 2 + h2 [\u2212xR,\u2212yR, h] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002681_j.jmatprotec.2016.05.029-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002681_j.jmatprotec.2016.05.029-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of experimental setup for fluoroscopic observation during laser cutting",
        "texts": [
            " Specimen dimensions were 4 \u00d7 120 \u00d7 4 mm (Fig. 1), set in consideration of the limited penetrative power of the X-rays. Two holes of 2.2 mm in diameter were drilled, one on each end of the specimen, for fastening to a jig. Another hole, this one also of 2 mm diameter, was drilled (not pierced) to provide a starting point for the laser cut. After surface degreasing with ethanol, the specimen was cut by laser in the longitudinal direction. A schematic diagram of the experimental setup is shown in Fig. 2. Laser cutting equipment was used together with fluoroscopic observation equipment. The laser cutting equipment consisted of a 16-kW disk laser (wavelength: 1030 nm; beam parameter product: 8 mm\u2219mrad; fiber diameter: 200 \u03bcm) and a cutting head. Both were arranged perpendicularly to the longitudinal axis of the specimen. The fluoroscopic observation equipment consisted of a micro-focus X-ray generator and an X-ray image intensifier, both of which are components of a high-brightness four-dimensional X-ray transmission system for the visualization of welding and joining processes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002005_toh.2014.2330300-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002005_toh.2014.2330300-Figure4-1.png",
        "caption": "Fig. 4. Definition of variables for the rendering. d is exaggerated for a clear view.",
        "texts": [
            " After again constructing vector triples, (d, _d, fT ), three Hunt-Crossley models are identified for the three Cartesian directions, denoted by Gx\u00f0dx; _dx\u00de, Gy\u00f0dy; _dy), and Gz\u00f0dz; _dz\u00de. The complete identification process takes less than a few minutes. Note that no geometric model is required for our approach. During rendering, a user first makes contact with the breast mock-up (i.e., the no-inclusion model), initiating the virtual inclusion rendering. To detect the initial contact, force sensor readings are employed (see [3] for more details). After contact, a user applies forces, possibly at multiple contacts, and causes the mock-up to deform, as shown in Fig. 4. The applied forces displace the contact point on the surface from the initial position pHs; to the current haptic point pH; \u00f0t\u00de. In addition, the inclusion is displaced from pT0 to pT \u00f0t\u00de. The force causing the inclusion to move is proportional to the inclusion response at the user\u2019s hand fT; \u00f0t\u00de in (1). Note that the mutual effects between the contact points due to the inclusion will be based on inclusion movement estimation. This will be described further below. Thus, fT; \u00f0t\u00de is directed from the inclusion position to the tool tip, with a magnitude of fT; \u00f0t\u00de such that fT; \u00f0t\u00de \u00bc fT; \u00f0t\u00de pH; \u00f0t\u00de pT \u00f0t\u00de jpH; \u00f0t\u00de pT \u00f0t\u00dej : (4) In (4), the unknown values, fT; \u00f0t\u00de and pT \u00f0t\u00de, need to be estimated during the rendering"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001864_20140313-3-in-3024.00082-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001864_20140313-3-in-3024.00082-Figure3-1.png",
        "caption": "Fig. 3. Twin Rotor MIMO System Laboratory Setup",
        "texts": [
            " Although the MIMO system control problem posed by the the TRMS setup and a real helicopter is similar, the major difference lies in the fact that in a helicopter, the aerodynamic force is controlled by varying the angle of attack of the rotors, while for a TRMS setup, the angle of attack is fixed with the aerodynamic force being controlled by varying the speed of the rotors. The highly nonlinear cross-coupled rotor dynamics, along with inaccessibility to many of the system states for measurement makes controlling the twin-rotor MIMO system a challenging 978-3-902823-60-1 \u00a9 2014 IFAC 740 10.3182/20140313-3-IN-3024.00082 control problem. The control objective is to make the beam of the TRMS move quickly and accurately to desired vertical and horizontal attitudes. The TRMS setup, as shown in Figure 3, consists of two rotor propellers perpendicular to each other and joined by a beam pivoted on its base that can rotate freely in the horizontal and vertical planes. The two rotors are driven by DC motors and are located at each end of the beam. The beam movement is controlled by changing the voltages to the DC motors, which in turn controls the angular velocities of the two rotors.The system is equipped with a counterweight pendulum hanging from the beam which dampens the TRMS motion and helps balance the angular momentum of the system under steady-state, or fixed loading conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000754_j.petlm.2015.10.012-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000754_j.petlm.2015.10.012-Figure2-1.png",
        "caption": "Fig. 2. Equivalent stress of the hollow c",
        "texts": [
            " The 140 kN bit pressure is selected to analyze the working performance in the paper. Exterior loading is applied on the roller teeth, but to ensure the convergence of the results, the single force of teeth is defined as the surface load acting on the cutting edge. At the same time, the friction torque is applied on the teeth. Placement constraint is fixed on the large end face of claw journal, and the symmetry constraint is fixed on the section of the model. Equivalent stress on the hollow cylindrical roller bearing of cone bit is shown as Fig. 2. The stress on the cone mainly focuses on the double ends of the large diameter hole and locking filling slot of cone. The stress on the claw journal is mainly concentrated on the thrust surface, cone transition chamfering, ends of the big journal, left side of the small journal and the transition fillet of the thrust surface and small journal surface. The stress of the hollow roller is mainly concentrated near the contact line, and the stress on the left is greater than that on the right side, which indicates that rollers under the actual working conditions bear the partial load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002740_b17506-6-Figure5.61-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002740_b17506-6-Figure5.61-1.png",
        "caption": "FIGURE 5.61 Voltage-limiting ellipses for IPM machines. (a) Voltage limiting ellipses and (b) voltagelimiting ellipses for different speeds.",
        "texts": [
            "57 and neglecting the voltage drop on the stator resistance: ( ) ( ) max\u2212 + + \u2264 \u21d2 \u239b \u239d\u239c \u239e \u23a0\u239f + + \u239b \u239d \u03c9 \u03c9 \u03c9\u03c8 \u03c8 L I L I V I L L I L q sq d sd m s sq d q sd m d 2 2 2 2 / \u239c \u239e \u23a0\u239f \u2264 \u239b \u239d\u239c \u239e \u23a0\u239f \u21d2 \u239b \u239d\u239c \u239e \u23a0\u239f + +\u239b \u239d\u239c \u239e \u23a0\u239f \u2264 2 2 2 2 1 V L I a I L b s d sq sd m d max \u03c9 \u03c8 (5.58) Id (A) Iq (A) L d i ncre asi ng L d de cre asi ng 0 Ld, Lq +25% Ld +50% Ld \u201325% Ld \u201350% Ld FIGURE 5.60 Sensitivity of MTPA curve for variations in Ld. D ow nl oa de d by [ C or ne ll U ni ve rs ity ] at 1 0: 56 0 3 A ug us t 2 01 6 where a V L b V L s q s d = =max max \u03c9 \u03c9 (5.59) For the given Vsmax and \u03c9, Equation 5.58 represents an ellipse as shown in Figure 5.61a. The center of the ellipse is located at (\u2212(\u03c8 m/Ld), 0), which is the characteristic current of the machine. The eccentricity of the ellipse can be represented as e b a b e V L V L V L L L s d s q s d d q = \u2212 = \u2212 = \u2212 \u239b \u239d\u239c \u239e \u23a0\u239f 2 2 2 2 2 1 ( ) ( )max max max \u03c9 \u03c9 \u03c9 (5.60) As shown in Figure 5.61b, the ellipse shrinks inversely with the rotor speed. The shape of the ellipse depends on the saliency ratio. In addition, the d-and q-axis currents must satisfy: i i Isd sq s 2 2 2+ \u2264 max (5.61) Equation 5.61 represents a current circle centered at the origin with a radius of Is max as shown in Figure 5.62a. Unlike the voltage-limiting ellipses, current-limiting circles remain constant for any speed. Since both Equations 5.57 and 5.61 should be satisfied during the operation, for a given rotor D ow nl oa de d by [ C or ne ll U ni ve rs ity ] at 1 0: 56 0 3 A ug us t 2 01 6 speed, the current vector can be located anywhere inside or on the boundary of the overlap area between the voltage-limiting ellipse and current-limiting circle as shown in Figure 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002583_icma.2014.6885682-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002583_icma.2014.6885682-Figure3-1.png",
        "caption": "Fig. 3. Link configuration of the robot",
        "texts": [
            " The position of joint j with respect to the origin of the world coordinate frame can be described by r\u3008j\u3009/0, while its axis orientation by the unit vector aj and its joint variable by \u03b8j . The link j has a mass distribution characterized by the mass mj , the position of its center of mass described by rcom\u3008j\u3009/0 and the tensor of inercia with respect to this point (and described in the world coordinate frame) Icom,j . Additionally, the joint j also drives all the link structure connected to the link j. Let us denote as m\u0303j , r\u0303com\u3008j\u3009/0 and I\u0303com,j the mass, the position of the center of mass and the inertia tensor of all this link structure driven by the joint j. See Fig. 3. Then, the additional linear and angular momentum with respect to the origin of the world coordinate frame for the robot which are yielded by \u03b8\u0307j (Pj,0 and Lj,0 respectively) can be calculated as Pj,0 = m\u0303jaj \u03b8\u0307j \u00d7 ( r\u0303com\u3008j\u3009/0 \u2212 r\u3008j\u3009/0 ) = Mj,0\u03b8\u0307j (27) Lj,0 = r\u0303com\u3008j\u3009/0 \u00d7 Pj,0 + I\u0303com,jaj \u03b8\u0307j = ( r\u0303com\u3008j\u3009/0 \u00d7Mj,0 + I\u0303com,jaj ) \u03b8\u0307j = Hj,0\u03b8\u0307j (28) Then, M\u03b8\u0307,0 = [ M1,0 M2,0 \u00b7 \u00b7 \u00b7 Mn,0 ] (29) H\u03b8\u0307,0 = [ H1,0 H2,0 \u00b7 \u00b7 \u00b7 Hn,0 ] (30) H\u03b8\u0307,com = H\u03b8\u0307,0 \u2212 r\u0302com/0M\u03b8\u0307,0 (31) Finally, let us consider (10) and substitute (31) into it in order to get a final expression for H\u03b8\u0307,zmp: H\u03b8\u0307,zmp = r\u0302com/zmpM\u03b8\u0307,com + ( H\u03b8\u0307,0 \u2212 r\u0302com/0M\u03b8\u0307,0 ) = H\u03b8\u0307,0 \u2212 r\u0302zmp/0M\u03b8\u0307,0 (32) Given that M\u03b8\u0307,0 = M\u03b8\u0307,com as Pj,0 = Pj,com"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000199_s00170-014-6551-7-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000199_s00170-014-6551-7-Figure1-1.png",
        "caption": "Fig. 1 LT kinematics",
        "texts": [
            " One beam is used as a reference, and the other beam is reflected back from a retro-reflector. The laser beam hits the retro-reflective target off-centre, and it is then reflected back parallel to the incident beam. A sensor measures the displacement. The LT measures the position of the target in spherical coordinates (d: radial distance; \u03b8: azimuth angle; \u03c6: elevation angle). The interferometer measures distance, d, and the horizontal and vertical encoders measure the azimuth and elevation angles, \u03b8 and \u03c6, respectively, as shown in Fig. 1. This section presents twomathematical models of a LT having the laser source in the rotating head. 3.1 Geometric error model The geometric error model of LTs having the beam source in the rotating head was presented in Loser and Kyle [19] and studied in Muralikrishnan et al. [16] and Hughes et al. [30]. This model is based on physical alignment deviations, and calibration corrections are based on 15 parameters, representing the influence of a particular geometric error in the overall error of the equipment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001907_tmag.2015.2448519-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001907_tmag.2015.2448519-Figure3-1.png",
        "caption": "Fig. 3. Current models for (a) cuboidal and (b) cylindrical permanent magnets.",
        "texts": [
            " The velocity oscillates with a frequency of \u03c9v = 2\u03c0 fv around a nominal value of v0 with an amplitude of v1 v(t) = v0 + v1 cos \u03c9v t . (11) Hence, the time-dependent position of the magnetic field source is given by \u03be(t) = v0t + v1 \u03c9v sin \u03c9v t . (12) The ratio v1/\u03c9v can be interpreted as a displacement amplitude of the magnetic field source. In the present context, rectangular and circular coil shapes are considered. In the further course of this paper, the planar cases are extended to stacked configurations as they are shown in Fig. 3. This is equivalent to the current model of permanent magnets [39]. Hence, cuboidal and cylindrical permanent magnets are included in the present analysis. 1) Rectangular Current Sources: The x- and y-components of the source current of a rectangular coil are given by I (rect) x = Is [ rect ( x \u2212 \u03be(t) a ) \u03b4 ( y + b 2 ) \u2212 rect ( x \u2212 \u03be(t) a ) \u03b4 ( y \u2212 b 2 )] (13a) I (rect) y = Is [ rect ( y b ) \u03b4 ( x \u2212 a 2 \u2212 \u03be(t) ) \u2212 rect ( y b ) \u03b4 ( x + a 2 \u2212 \u03be(t) )] (13b) where rect(\u00b7) is the rectangular function, \u03b4(\u00b7) is the Dirac delta distribution, and Is is the current strength",
            " The infinite bandwidth of the function leads to a sum of harmonics with their amplitudes determined by the velocity oscillation amplitude v1, the frequency \u03c9v and kx . In numerical calculations, the sum in (30) is truncated after N terms. Finally, the function Sc,m(k) includes the height information of the magnetic field source. The index w and m distinguishes between the single wire with c = 0 or a magnet of height c > 0, respectively. The solution of a single wire can be extended to permanent magnets by means of the surface current model shown in Fig. 3. It is assumed that the permanent magnets have a height of c and their lower surface is located at a liftoff distance h above the plate. The total force is determined by integrating the force on each wire over the z coordinate, resulting from the superposition of the total distorted magnetic flux density generated from all wires. Hence, (28) becomes Fx |z(t) = 1 \u03c02 \u222b \u221e \u2212\u221e \u222b \u221e \u2212\u221e Ix |zG(t, k) f (t)\u22121 \u00d7 \u222b h+c h \u222b h+c h e\u2212k(z+z\u2032) dz dz\u2032 \ufe38 \ufe37\ufe37 \ufe38 Sm(k) dkx dky . (32) Carrying out the integration over z and z\u2032, one gets the expression of the function Sm(k) valid as long as c > 0 Sm(k) = 1 k2 ( e\u2212kh \u2212 e\u2212k(h+c))2 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001477_1.5025605-Figure38-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001477_1.5025605-Figure38-1.png",
        "caption": "FIG. 38. The rate of fluid squeeze-out for the sphere with radius R is the same as for a circular disc with radius r = (2Ru)1/2.",
        "texts": [
            " Thus we assume that the attractive glass-rubber interaction matters only in that it determined the initial (average) surface separation u(0) = u0 when F = 0. When a rigid ball (radius R) is squeezed against a flat rigid surface in a fluid with the viscosity \u03b7, the relation between the force F and the (minimum) surface separation u is given by75,76 F = \u22126\u03c0\u03b7R2 u\u0307 u . (I1) As shown in Appendix J, this equation can also be derived by considering the fluid squeeze-out for a circular plate with radius r = (2Ru)1/2 (see Fig. 38). The interfacial separation velocity u\u0307(t) is not the same as the drive velocity vz because of the elasticity in the system. Thus the nylon rope has an effective spring constant k \u2248 100 \u2013200 N/m (see Appendix H), and the rubber substrate too will deform elastically in response to the viscous force F acting on it. As a result F = k\u2217(s \u2212 u), where s = s0 + vt is the drive displacement and k\u2217 < k is an effective (combined) spring constant (see Appendix J). Combining this equation with (I1) gives the equation of motion for u(t) (here we have neglected inertia effects) 6\u03c0\u03b7R2 u\u0307 u = k\u2217(s \u2212 u)",
            " VI E, where we found that for the clean glass ball the longer contact time results in a stronger adhesive interaction between the glass ball and the rubber, and hence to a smaller u0. Consider the fluid squeeze flow between a rigid sphere and a flat rigid surface. Let u(t) denote the minimum separation between the surfaces. Most of the resistance toward fluid squeeze-out occurs in the region where the separation between the surfaces arises from the area where the separation between the surfaces is of order a few times u(t), say when the separation varies from u(t) to 2u(t). When u R, the radius of this circular region is r \u2248 (2Ru)1/2 (see Fig. 38). In a first approximation, we can replace the sphere with a circular disc with radius r separated from the flat substrate with the distance u. For a circular disk, the relation between the applied force F and the separation u is given by the well-known relation (see, e.g., Refs. 1 and 2) F = \u2212 3\u03c0 2 \u03b7r4 u\u0307 u3 . Substituting r \u2248 (2Ru)1/2 in this equation gives F = \u22126\u03c0\u03b7R2 u\u0307 u , (J1) which agrees with the exact result (I1). Let us now study the fluid squeeze flow when the drive is moved away (pull-off) from the substrate with the speed v "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002948_icems.2019.8922014-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002948_icems.2019.8922014-Figure1-1.png",
        "caption": "Fig. 1. Photograph of Proposed SRM",
        "texts": [
            " Appling the value to the simulation, as a results, it is revealed that the motor efficiency improves thanks to the reduction of eddy current loss and windage loss by the proposed structure. Finally, the number of division of the rotor core is optimized at each driving points. As a result, it is found that 10 division of the rotor core is the best at low speed and low to middle torque area. On the other hand, 30 division of the rotor core is the best at high speed and low torque area. II. PROPOSED SRM In this study, the SRM is proposed as an alternative to PMSM that has already been adopted as the traction motor. Fig. 1 shows the proposed SRM (3-phase, 8-poles, and 12- slots). As described in Chapter I, mechanical strength of the SRMs is higher than PMSMs. Because of this advantage, the SRM can obtain high output power by high speed drive. TABLE I shows the specifications of the proposed SRM and the conventional PMSM. As shown in TABLE I, maximum speed of the proposed SRM is three times higher than the conventional PMSM. As a result, the volume of the proposed SRM is reduced by 53% compared with the conventional PMSM at the same maximum output power"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001324_j.triboint.2018.01.017-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001324_j.triboint.2018.01.017-Figure6-1.png",
        "caption": "Fig. 6. Structure of the spiral groove gas face seal in the experimental chamber.",
        "texts": [
            " 4, consists of an electric motor, two couplings, a torque-speed sensor, two bearings, an experimental chamber, an air supply system, etc. The infinitely variable control motor can be adjusted from 0 to 3000 r/min. The torque-speed sensor with 10-Hz sample frequency can achieve the real-time value about the speed, torque and power. The air supply system can provide the experimental chamber the dry air ranging from 0 to 0.8MPa. The structure of the experimental chamber [31,32] is displayed in Fig. 5. A symmetric double spiral groove gas face seal produced by John Crane Inc. is adopted, as shown in Fig. 6. The rotor is located between the two flexibly mounted stators, thus the two stators can track the rotor in the axial and angular modes. The present study only involves the stator that is close to the end cover of the chamber. When the dry air is inputted into the chamber: on the one hand, it passes through the rotor and the stator that is far away from the end cover, and then leaves the chamber via the bearing box; on the other, it, successively, passes through the rotor and the stator that is close to the end cover, a hole on the chamber shell, a short pipe, and a digital flowmeter with 10-Hz sample frequency"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000600_0954406213477777-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000600_0954406213477777-Figure5-1.png",
        "caption": "Figure 5. A range of of bevel gear with: (a) chipped tooth, (b) missing tooth and (c) worn gear.",
        "texts": [
            " This could be used for the simulation of a range of machine faults like in the gearbox, shaft misalignments, rolling bearing damages, resonances, reciprocating mechanism effects, motor faults, pump faults, etc. In the MFS experimental setup, three-phase induction motor is mounted to the rotor that is connected to the gear box through a pulley and belt mechanism. The gear box and its assembly are illustrated in Figure 4. In the study of faults in gears, three different types of faulty pinion gears namely the chipped tooth (CT), missing tooth (MT) and worn tooth (WT) along with normal gear (or no defect i.e. ND) were used (illustrated in Figure 5). The real time data in time domain were measured using a tri-axial accelerometer (sensitivity: x-axis 100.3mV/g, y-axis 100.7mV/g, z-axis 101.4mV/g) mounted on the top of the gearbox (illustrated in Figure 6) and the data acquisition hardware. Measurements were taken for the rotational speed of 10 to 30Hz in intervals of 2.5Hz for each of four conditions. For each measurement set, 300 Table 2. Optimization fitness function and design parameters. Fitness function Design parameters Bounds C-SVC with RBF kernel Maximize f(x)\u00bc (number of correctly predicted data/total number of testing data) 100% X \u00bc C T For : 0\u20131 For C: 0\u20131"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003765_iros.2018.8594316-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003765_iros.2018.8594316-Figure3-1.png",
        "caption": "Fig. 3: The size of newly developed hand.",
        "texts": [
            " Concerning MP joints, it is difficult to arrange actuators to control muscles independently because there are eight interossei muscles to move four fingers in quite a narrow space. Therefore, we made a novel thumb CM joint by the combination of three machined springs and realized the joint with DOFs including a joint with wide range of motion. And we applied a variable rigidity mechanism of agonistic tendon drive and placed one actuator to move this mechanism toward both of the fit of finger when grasping and the transmission of force with the side of fingers. The whole view and size of the hand we developed in this study are shown in Fig. 3. It is a five-fingered hand imitating human hands. It is connected to forearm through the wrist joint and it is moved by the muscles placed in forearm. The fingers of this hand is shown in Fig. 4. It has finger joints made of machined springs as with the hand of \u201cKengoro\u201d. The machined spring has flexibility and toughness because it is a kind of springs made of metal. In addition, it has an advantage that it can be connected firmly to other parts because it can be designed integrally with attachments"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003816_s12541-019-00014-2-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003816_s12541-019-00014-2-Figure1-1.png",
        "caption": "Fig. 1 Blade bearing of the controllable pitch propeller",
        "texts": [
            " The problem description of the assembly deviation under actual working conditions is presented in Sect.\u00a02. The actual working conditions are introduced in Sect.\u00a03. The modeling of the assembly deviation is discussed in Sect.\u00a04. An application is presented in Sect.\u00a05 to illustrate the approach of this paper, and Sect.\u00a06 provides the discussion and conclusion. In this paper, the blade bearing of a CPP is regarded as an application to address the methodology of assembly deviation with considering the dynamic actual working condition. Figure\u00a01 (left) displays the blade bearing of a CPP, and Fig.\u00a01 (right) presents the partial enlarged view of an O-ring hydraulic seal. This assembly consists of blade foot, blade carrier, O-ring and hub. The blade carrier is mounted on the hole of the hub. The O-ring is extruded to fill up the O-groove between the blade foot and hub; and it can form O-ring hydraulic seal to prevent the seawater flow into the hub and the hydraulic oil flow into the seawater. Simultaneously, the blade carrier and bladed foot are fastened with bolts, and there is no relative displacement between them. During the actual working conditions, the deformations of the parts of the blade bearing are affected by the centrifugal load, hydrodynamic load, and gravity. Besides, the temperature of the oil in the hub can also lead to deformations, as shown in Fig.\u00a01. Moreover, during the service process, the fluctuations of the sea water could result in the fretting behavior of the blade bearing in a CPP. In fact, when the wear of contact surface 3 increases, the gap (contact surface 1 in Fig.\u00a01) between the blade foot and hub can also increase, and it will lead to a rebound of the O-ring. When the decreased compression ratio is less than the required compression ratio, leakage will occur. In engineering design, if the initial gap of contact surface 1 is designed too large, then with the occurrence of these deformations and wear, its service life would be decreased. In this paper, the service life refers to the seal failure time of the blade bearing when its assemble gap increases to spill, and here, the fatigue life of O-ring and blade bearing is not considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000796_s1068371215070111-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000796_s1068371215070111-Figure2-1.png",
        "caption": "Fig. 2. Transformed graph of the investigated EMS with EC.",
        "texts": [
            "c(S)FCR(S)Fconverter(S)Fe(S)Fengine(S)Ffv(S) and F02(S) = FVR(S)Fconverter(S)Fe(S)Ffc(S) are the trans missions of the graph loops taking into account condi tionally opened circuits of regulations of velocity (VRC) and current (CRC), loop F03(S) = Fengine(S)Fe(S) takes into account the influence of the counter EMF of the engine, and loops F04(S) = Felastic(S)Fengine(S) and F05(S) = Felastic(S)Fm(S) take into account the inertia and elastic\u2013dissipative prop erties of the SMP. To solve the problem of SRS synthesis permitting an EMS with EC minimizing the oscillations of the torque in the elastic element Melastic(S) upon a change of disturbing action Md(S), the initial graph of the EMS with EC in Fig. 1 is transformed to the form shown in (Fig. 2). In Figs. 1 and 2, to reduce designa tions of the junctions and transmission of the graph branches, they were given without any signs of opera tors (S) designating the application of Laplace trans formation. Let us determine the operator expression of the torque in the elastic element based on the graph in Fig. 2a: (2) where Fc.c(S) = F02(S)/{F01(S)[1 + F02(S)]} and Fc.s(S) = F01(S)/{Fbv(S)[1 + F01(S)]} are transmissions of the graph branches defining the properties of the closed CRC and VRC. To simplify the investigations. let us transform the graph in Fig. 2a to the form shown in Fig. 2b, where branch Felastic.m(S) = (TdS + 1)/(TsS){1 + [(TdS + 1)/(TsTmS2]}\u20131 characterizes the properties of the SMP and branch FED(S) = Fc.s(S)[FVR(S)Ff.c(S)Fc.c(S)]\u20131 takes into account the properties of the electric drive with the SRS. Usually, during synthesis of an ED SRS with abso lutely rigid transmissions, the equivalent properties of the CRC and VRC are described by aperiodic links of the first order, which, during research on the EMS with EC, can lead to significant research errors. In this research, depending on the structure and parameters of SRS regulators, it is suggested to consider the dynamic properties of the electric drive by transmis sions of the second and third orders",
            " 7 2015 PYATIBRATOV Research on the efficiency of the electric drive damping properties shall be carried out in the terms of the EMS with EC having the following parameters of the invariable part of the system: Tconverter = 0.005, \u03c4 = 0.003 s, kelastic = 8.2, Telastic = 0.132 s, Ts = 0.0134 s, Td = 0.005 s, Tm = 0.38 s, kconverter = kengine = km = kfc = kfv = 1, and Tfc = Tfv = 0. To determine the conditions providing the best damping of torque oscillations in the elastic element Melastic.ext by the electric drive, let us consider frequency characteristics of the transmission of branches Felast.m and FED of the graph given in Fig. 2b. In Fig. 3, LACH of branch Felast.m taking into account the SMP properties, LACH LED of the back branch taking into account the ED, and LACH Lr showing change of torque in the elastic element Melastic.ext are given. Index numbers i = 1\u20134 of characteristics LED and Lr correspond to the variants of implementa tion of the SRS represented in Table 2. LACH Lr.t. shows the change of the torque of the elastic element Melastic.ext in the SMP. The research showed that, in this case, the torque amplitude in the elastic element at the resonance frequency Melastic"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002336_1.g000332-Figure18-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002336_1.g000332-Figure18-1.png",
        "caption": "Fig. 18 Geometry of cannon\u2013-radar\u2013projectile problem.",
        "texts": [
            " For these reasons, using an EKF or SDRE filter with states x _x x T is not suggested for measuring the ballistic coefficient \u03b2. C. Estimating the Trajectory of a Cannonball In the previous two examples, we have considered cases where the system dynamics is nonlinear and the measurement equations are linear. In this example, we consider a casewhere the system dynamics is linear and the measurement equations are nonlinear. In light of ([14] pp. 331\u2013349, Chap. 9), we consider a radar tracking a cannonlaunched projectile, or cannonball, in two-dimensional Cartesian space. Figure 18 lays out the problem geometry. We assume that the cannon is located at the origin, a radar is located at xR; yR , the projectile is located at xT t ; yT t at time t, and the radar is tracking range r and angle \u03b8 between itself and the projectile. The state we use for the tracking filters is xT _xT yT _yT T . One can show that, under the influence of gravity g, the system dynamics can bewritten as d dt 2 664 xT _xT yT _yT 3 775 2 664 0 1 0 0 0 0 0 0 0 0 0 1 0 0 0 0 3 775 2 664 xT _xT yT _yT 3 775 2 664 0 0 0 \u2212g 3 775 2 664 0 0 1 0 0 0 0 1 3 775 w _xT w _yT y \u03b8 r tan\u22121 yT \u2212 yR\u2215xT \u2212 xR xT \u2212 xR 2 yT \u2212 yR 2 p v (25) We first go about setting up an SDRE factorization for the observation equation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003182_j.compag.2017.08.016-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003182_j.compag.2017.08.016-Figure1-1.png",
        "caption": "Fig. 1. The structure diagram of operation boat (1. the hull; 2. the cutting device for aquatic weeds; 3. the conveying device for aquatic weeds; 4. the collection box for aquatic weeds; 5. the aquatic weeds paving device; 6. the paddle wheel; 7. the operation station; 8. the feeding machine.)",
        "texts": [
            " Section 2 presents the general structure of the agriculture operation boat (AOB). Section 3 presents the dynamics of the path-following errors. In Section 4, we design the proposed RMPC. Simulations and experiments are described in Sections 5 and 6, respectively. At last, Section 7 makes conclusions of the whole paper. 2. General structure of the agricultural operation boat The AOB mainly consists of the hull, the aquatic weed clearing device, the feeding machine, and the paddle wheel driving device, whose structure diagram and main parameters are shown in Fig. 1 and Table 1, respectively. In Fig. 1, the cutting and conveying devices placed on the bow can cut and collect aquatic weeds, the collection box in the middle of the hull can store cut weeds; the paving device behind the collection box can pave aquatic weeds and avoid the accumulation, and the feeding machine in the stern can cast baits. The paddle wheels equipped on both sides of the hull is the driving device, which can avoid weeds winding. Electricity of the whole system is supplied by a 48 V lithium battery with 120AH capacity, which has the advantages of no pollution, high efficiency, and low noises"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001808_1.4037667-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001808_1.4037667-Figure2-1.png",
        "caption": "Fig. 2 Rotor-bearing axial coordinate configuration",
        "texts": [
            " x is the angular velocity of the rotor. l is the air viscosity \u00f0l \u00bc 18:6 10 6 Pa s\u00de and Rg is the gas constant for the air \u00f0Rg \u00bc 287:05 J=\u00f0kg K\u00de\u00de. T is the temperature of the supplied air (T\u00bc 293 K), and _ms is the air mass flow rate through the orifice tube and it is calculated from isentropic flow model through the orifice [17]. The boundary pressure on the both ends of the bearing is the atmospheric pressure \u00f0Pa \u00bc 101 kPa\u00de. zj\u00f0j \u00bc 1; 2\u00de represents the bearing local coordinate along the axial direction. Figure 2 depicts the relation between the global axial coordinates \u00f0Zj\u00de for the locations of the bearings and local coordinate zj. The film thickness for each bearing (j) is expressed by motion variables of the rotor center of gravity, axial location of the bearings, and other parameters such as hj\u00f0h; zj\u00de \u00bc C\u00fe \u00f0X \u00fe wZj\u00decos h\u00fe \u00f0Y nZj\u00desin h \u00fewj\u00f0h; zj\u00de rp cos\u00f0h hp\u00de (3) 012506-2 / Vol. 140, JANUARY 2018 Transactions of the ASME Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/ on 10/04/2017 Terms of Use: http://www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003973_icieam.2019.8743076-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003973_icieam.2019.8743076-Figure6-1.png",
        "caption": "Fig. 6. Temperature distribution inside the induction motor, at t = 8000 s, considering a faulty motor condition. The short-circuit resistance is set at 0.001 .",
        "texts": [
            " Due to the large data sets produced while running the thermal simulations, the analysis of the simulation results will solely focus on the analysis of the temperature distribution inside the motor at the instants in which the thermal steady state is reached. According to IEEE and IEC standards, the thermal steady state of electrical machines is reached when the temperature increment along one hour is less than 2 \u00b0C [11], [12]. In this context, an analysis based on the observation and relative comparison of scenarios will be employed. Fig. 3 to Fig. 6 depict the temperature distribution inside the motor, in \u00b0C, considering the motor operation under healthy and faulty conditions. Note that the temperature distributions presented in Fig. 3 to Fig. 6 refer to the instant of time at which the thermal steady state condition was already met. In the conditions of load torque and initial room temperature defined for the simulations, the thermal steady state conditions are observed before the instant t = 8000 seconds. As depicted in Fig. 3, there is a consistent and uniform distribution of temperatures inside the motor, typical of a healthy operation scenario. As expected, the rotor is the warmer part of the motor, followed by the stator. The rotor temperature is indeed quite uniform",
            " As a result, the progressive decrement of temperature in the direction of the stator outer boundary disappears. In this case, a simple observation of the stator surface temperature would allow to clearly notice the thermal unbalance along the stator. Despite the severe increment of temperature occurring near the shortcircuited turns, the temperature of the warmest regions remains slightly below the critical temperature, at which the integrity of the windings\u2019 insulation is sustained (155 \u00b0C). Fig. 6 depicts the temperature distribution map inside the motor for the scenario of a pure short-circuit between 6 turns of phase W. As witnessed in Fig. 6, the condition with lowest shortcircuit resistance leads to a massive overheating of the windings located nearby the short-circuited turns. Even though the excessive heating is pretty much concentrated in space, such overheating also spreads to the rest of the stator and to the rotor, leading to a non-uniform heating of the motor. Since the insulation class of the motor windings is class F, the motor would fail well before verifying such a violent short-circuit condition. Just like in the previous scenario (refer to Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000243_1.4798508-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000243_1.4798508-Figure1-1.png",
        "caption": "FIG. 1. Two ways for compensating gravity: a) the orientation of the microrobot is arbitrary (e.g., coinciding with the direction of motion\u2013\u2013gradientbased); b) the orientation of the microrobot is constrained (i.e., coinciding with that of the magnetic force\u2013\u2013orientation-based).",
        "texts": [
            " In the following, we use the parameter dS \u00bc Dq=jDqj defining the direction of gravitational shifting (positive: upward; negative: downward). All the forces in Eq. (1), together with the velocity and the gravity vectors v * and g * , lie in the plane identified by v * and the z axis of the global reference system. Hence, we define the angles h and hF that represent, respectively, the direction of motion of the microrobot and the orientation of the magnetic force in the v * z-plane with respect to the xyplane (see Figure 1). In the following, we also refer to the angle a, defined as the orientation of the v * z-plane with respect to the xz-plane. The amplitude of the magnetic force FM required for reliably propelling the microrobot within the fluid can be obtained by FM \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F2 D \u00fe F2 G 2dSFDFG sin h q \u00bc VrB M; (2) where FD and FG are the amplitudes of the drag and gravitational force, respectively, rB is the norm of the magnetic gradient tensor (representing its amplitude), and M is the amplitude of the magnetization of the microrobot",
            " Here the symbol indicates the unit vector, fD \u00bc FD=FM, and fG \u00bc FG=FM. C is a matrix, defined as the ratio between the magnetic gradient tensor and its norm rB, describing the influence of the magnetic gradient tensor on the orientation of the resulting magnetic force. The angles hM and aM identify the orientation of the magnetization vector M * . For better understanding the practical consequences of the general relation given in Eq. (4), here we consider two specific and opposite strategies for orienting the magnetic force along the required direction hF (see Figure 1). One method enables the compensation of gravity while also allowing the magnetization vector, and thus the microrobot, to have an arbitrary orientation, e.g., to be aligned to the direction of motion, such that aM \u00bc a and hM \u00bc h 6\u00bc hF (see Figure 1(a)). Therefore, the microrobot moves as an ideal, not-sinking micro-swimmer. This strategy is based on tailoring the magnetic gradient on purpose (gradient-based gravity compensation). Jeong et al.14 implement in their magnetic navigation system a particular embodiment of this method, in which C is defined as a diagonal matrix. Such choice is commanded by the standard architecture, i.e., an assembly of Helmholtz and Maxwell coils, of the adopted system. The effect of the singularities inherent in this definition of C is avoided by moving the microrobot on a test bed that contributes to the compensation of gravity",
            " This is the case of both the OctoMag16 and MiniMag20 systems, where complex magnetic gradients are generated and 5-degrees-of-freedom microrobotic navigation is enabled. However, both the development and control of such advanced systems are ways more complex than those of a standard one. The other method is based on constraining the orientation of the microrobot, which is aligned to the required magnetic force orientation, rather than to the direction of motion, such that aM \u00bc a but hM \u00bc hF 6\u00bc h (orientation-based gravity compensation, Figure 1(b)). In this case, the motion of the microrobot is not ideal, but the condition on the applied gradient reads C \u00bc I3 (being I3 the identity matrix). This method thus implies a simple and diagonal formulation of the magnetic gradient tensor. This can be obtained by orthogonal This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 128.59.226.54 On: Wed, 10 Dec 2014 16:04:51 assemblies of Maxwell coils, which are widely used, due to their simplicity, as a standard arrangement in magnetic navigation systems"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure4-1.png",
        "caption": "Fig. 4. Boundary fillet \u2014 type IIa of the rack-cutter.",
        "texts": [
            " Therefore curve \u03be allows us to define the following boundary condition: undercutting \u2014 type IIb is avoided if the real rack-cutter fillet AF is placed internally regarding the boundary fillet \u03be. In the case shown in Fig. 3b this condition is not satisfied and as a result the gear teeth are undercut \u2014 type IIb. 3. \u0415quations of boundary fillet curves 3.1. Boundary fillet \u2014 type II\u0430 (curve \u03b7) The equations of this curve are found using the theory of plane meshing [1], where one of the two conjugated profiles is set and the other one is obtained as an envelope of the relative positions which the specified profile occupies in the plane of the searched one. In this case (Fig. 4) the specified profile is the radial line l (axis OlYl) of the gear and the searched profile is the boundary fillet \u03b7 of the rack-cutter. In order to solve the problem of the geometrical synthesis, two mobile coordinate systems are introduced: XlOlYl is connected with the gear (the specified profile l); while X\u03b7O\u03b7Y\u03b7 is connected with the rack-cutter (the searched profile \u03b7). As the axis O\u03b7X\u03b7 (the centrode line) of the rack-cutter rolls without sliding on the reference circle (of a radius r) of the gear, the displacement s of X\u03b7O\u03b7Y\u03b7 is synchronized with the rotation of XlOlYl at an angle \u03c6, where s=r\u03c6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000787_j.ifacol.2015.10.272-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000787_j.ifacol.2015.10.272-Figure1-1.png",
        "caption": "Fig. 1. ROV operational configuration.",
        "texts": [
            " As an advantage, the proposed method permits to be applied in all thruster faulty situations. The equations describing the ROV dynamics have been obtained from classical mechanics (Conter et al., 1989; Longhi and Rossolini, 1989; Corradini et al., 2011a). The ROV considered as a rigid body can be fully described with six degrees of freedom, corresponding to the position and orientation with respect to a given coordinate system. Let us consider the inertial frame R (0, x, y, z) and the body reference frame Ra (0a, xa, ya, za) (Conter et al., 1989) shown in Fig. 1. Using the standard notation adopted by the SNAME (Society of Naval Architects and Marine Engineers), established in 1950, the ROV position with respect to R is expressed by the origin of the system while its orientation by the roll, pitch, and yaw angles (i.e., \u03c6, \u03b8, and \u03c8, respectively). Thus the dynamic equations of ROV motion in the body-fixed frame can be expressed by (Fossen, 2011) M\u03bd\u0307 +C(\u03bd)\u03bd +D(\u03bd)\u03bd + g(\u03b7) = \u03c4 (1) with M = inertia matrix (including added mass) C(\u03bd) = matrix of Coriolis and centripetal terms D(\u03bd) = drag matrix g(\u03b7) = vector of gravity, buoyancy forces and moments \u03c4 = control forces and moments acting on the ROV centre of mass where, \u03bd1 = [u v w] T = [x\u0307 y\u0307 z\u0307] T , \u03bd2 = [p q r] T =\ufffd \u03c6\u0307 \u03b8\u0307 \u03c8\u0307 \ufffdT , thus the ROV spatial velocity vector with respect to its body-fixed frame is \u03bd = \ufffd \u03bd1 \u03bd2 \ufffd (2) \u03b71 = [x y z] T , \u03b72 = [\u03c6 \u03b8 \u03c8] T , and thus the position and orientation state vector with respect to the inertial frame is \u03b7 = \ufffd \u03b71 \u03b72 \ufffd (3) The existing relationship between the velocity vector in the body-attached frame and the position vector in the inertial frame are given by \u03b7\u0307 = J(\u03b7)\u03bd (4) with J(\u03b7) = \ufffd J1(\u03b72) 03\u00d73 03\u00d73 J2(\u03b72) \ufffd (5) and J1(\u03b72) = \ufffd c\u03c8c\u03b8 \u2212s\u03c8c\u03c6 + c\u03c8s\u03b8s\u03c6 s\u03c8s\u03c6 + c\u03c8c\u03c6s\u03b8 s\u03c8c\u03b8 c\u03c8c\u03c6 + s\u03c6s\u03b8s\u03c8 \u2212c\u03c8s\u03c6 + s\u03b8s\u03c8c\u03c6 \u2212s\u03b8 c\u03b8s\u03c6 c\u03b8c\u03c6 \ufffd (6) J2(\u03b72) = \ufffd 1 s\u03c6t\u03b8 c\u03c6t\u03b8 0 c\u03b8 \u2212s\u03c6 0 s\u03c6/c\u03b8 c\u03c6/c\u03c6 \ufffd (7) where s(\u00b7) = sin(\u00b7), c(\u00b7) = cos(\u00b7) and t(\u00b7) = tan(\u00b7)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001777_978-1-4419-8420-3-Figure2.9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001777_978-1-4419-8420-3-Figure2.9-1.png",
        "caption": "Fig. 2.9 Inspection process on a simulated pipeline. a Upper simulated model. Lower regions guarded by different guard points encoded in different colors. b Damaged pipe with some holes. c Damages are detected, whose boundaries are extracted and shown in green",
        "texts": [
            " We apply our procedure on this platform using several complicated 3-D virtual pipelines. The simulated results are convincing and show the effective inspection on pipeline geometry. 32 2 Region-Guarding Problem in 3-D Areas If a hole appears on the pipeline, it can be identified online when the robot reaches the guarding point that covers this region and matches the captured range depth images with the stored templates. We simulate this on pipeline meshes M by randomly generating some missing regions. An experiment is shown in Fig. 2.9. A pipeline model and the necessary guards are shown in (a), where regions covered by each guard are rendered in a specific color for the visualization purpose. Any given region of the pipeline is covered by at least one guard and, therefore, is colorized. The height maps can then be generated as templates, measuring the \u201ccorrect\u201d distance from each guarding site to the pipe wall toward specific directions. This simulates range images obtained by a laser scanner. Now, we simulate the appearance of defected regions on the pipeline by generating some missing regions as shown in (b)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003114_lra.2017.2728200-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003114_lra.2017.2728200-Figure1-1.png",
        "caption": "Fig. 1: A quadruped robot navigating a complex 3D structure.",
        "texts": [
            " In environments that are unknown a priori or only partially known, a popular approach in legged robot navigation is to generate cyclic gaits and adapt gait parameters (e.g. stride length and stride height) in a reactive manner [3]. This has achieved significant success due to its simplicity and ease of implementation on real systems. However, this only succeeds in situations that have limited complexity and often fails in more challenging environments which require complicated full-body manoeuvres. In this paper, we present a full-body planning algorithm for legged robots capable of navigating such complex environments as shown in Fig. 1. Similar to earlier approaches [4] we decompose the problem of robot motion planning into body planning and planning of individual legs to enable stability and reachability. The use of the pre-computed Contact Dynamic Roadmap (CDRM) structure to identify footholds for individual legs is the crux of our approach. While previous work solves the problem sequentially, we explicitly determine the existence of valid footholds and stable configurations before accepting a body pose. This precludes conditions where the robot plans would have to be recomputed when the feasibility hypothesis fails",
            " This illustrates that the density of the CDRM is sufficient to exploit the full extent of the robot\u2019s dexterity, while still remaining quick to search. Full-body motion planning enables traversal of complex 3D structures, where cyclic gait-based planners would not succeed. Two sample scenarios shown in Figs. 1 and 9 show such complex scenarios which are composed of many individually challenging features, such as narrow passages and steps. We present the planning times and success rates for these two example problems in Table I, including a breakdown of each algorithm component. a) Truss: The scenario in Fig. 1 shows a robot climbing a complex 3D truss structure. The planned path is approximately 4m long and took 76.9s to plan with a success rate of 80%. As the planning time was approaching the limit of 120s, the success rate was not 100%. This occasional failure in complex environments is mitigated by the ability to re-plan due to the relatively short planning time. b) Pipe: A 7m path was planned through a pipe with several obstructions. This scenario is inspired by a robot with magnetic feet being used to perform an inspection task",
            " This took approximately 180s to generate a single valid sample, illustrating this approach is not feasible for planning with contacts as the probability of selecting valid contact configurations is small. 2377-3766 (c) 2017 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Next we ran the RB-PRM planner from the open-source Humanoid Path Planner package [23] on the same robot for two scenarios, the rubble (Fig. 10) and the truss (Fig. 1). The crux of the approach [4] is that if each robot leg is operating within its reachability limits, then it is assumed a stable fullbody configuration can be generated. For the rubble scenario shown in Fig. 10, our planner took on average 44s to generate a motion plan, while the RB-PRM planner took between 10s and 15s which was comparable to the 7s reported in [4] using the HyQ platform. We then tested both approaches on the truss shown in Fig. 1. Our planner successfully generated paths in 77s, but RB-PRM failed to plan the vertical transitions, likely as the reachability assumption did not hold. These results show that on planar terrains our planner is within the same order of magnitude performance as [4], while enabling planning in complex environments such as Fig. 1 where first computing a guide path and subsequently generating contacts may fail. We have presented CDRMs which extend DRMs with contact information. These can be used to rapidly identify footholds for legged motion planning. Pre-computing foothold and collision information ahead of time allows valid footholds for a body pose to be identified with minimal collision checks, by looking up this information directly from the CDRM. This allows using a sampling-based planner to generate full-body plans which satisfy leg contact and stability criteria constraints in complex 3D environments with high success rates"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003448_j.procir.2018.03.277-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003448_j.procir.2018.03.277-Figure5-1.png",
        "caption": "Fig. 5 Photograph of vise no. 3. Made by mechanical engineer students.",
        "texts": [
            " Findings There were three groups that chose the optional assignment with 3D-printing. One of the groups were design students, and two of the groups were mechanical engineer students. The students chose the group members themselves. 3.1. Practical results of the process The assignment was not carried out exactly as planned. The students were supposed to get the nylon parts before they started, but because the nylon printer had a failure and needed to be repaired, the students got the parts at the end of the project period. These are the white parts in Fig. 4 and Fig. 5. Nevertheless, all the groups managed to complete the assignment within the deadline, and ended up with a fully functional vise. The vices made by the mechanical engineer students are shown in Fig. 3 and Fig. 5. The vise made by the design students is shown in Fig. 4. The design students used this unforeseen opportunity to experiment and then print the missing parts in plastic with the cheap desktop printers, as shown in Fig. 6. They also managed to print the small screws in plastic, although steel screws were handed out (see the two screws at the bottom in Fig. 6). Beforehand it was considered unlikely that the threaded parts could be printed in plastic, but the design students did it anyway. They did however experience that the smallest plastic screws broke during assembly because they were not strong enough, and the threads on the spindle (see third part from the top in Fig",
            " Findings There were three groups that chose the optional assignment with 3D-printing. One of the groups were design students, and two of the groups were mechanical engineer students. The students chose the group members themselves. 3.1. Practical results of the process The assignment was not carried out exactly as planned. The students were supposed to get the nylon parts before they started, but because the nylon printer had a failure and needed to be repaired, the students got the parts at the end of the project period. These are the white parts in Fig. 4 and Fig. 5. Nevertheless, all the groups managed to complete the assignment within the deadline, and ended up with a fully functional vise. The vices made by the mechanical engineer students are shown in Fig. 3 and Fig. 5. The vise made by the design students is shown in Fig. 4. The design students used this unforeseen opportunity to experiment and then print the missing parts in plastic with the cheap desktop printers, as shown in Fig. 6. They also managed to print the small screws in plastic, although steel screws were handed out (see the two screws at the bottom in Fig. 6). Beforehand it was considered unlikely that the threaded parts could be printed in plastic, but the design students did it anyway. They did however experience that the smallest plastic screws broke during assembly because they were not strong enough, and the threads on the spindle (see third part from the top in Fig",
            " It is not possible to know how the quality of the drawings would have been if the students had not taken the optional additional work with 3D-printing. However, it should be noted that the 14 students participating had from none (0!) up to max three comments on their set of 2Ddrawings, which is an indication of an increased level of understanding when comparing with those students not taking on this task. Fig. 3 Photograph of vise no. 1. Made by mechanical engineer students. Fig. 4 Photograph of vise no. 2. Made by design students. Fig. 5 Photograph of vise no. 3. Made by mechanical engineer students. Fig. 6 The additional parts printed by the design students. Author name / Procedia CIRP 00 (2017) 000\u2013000 5 3.2. Cognitive result; students\u2019 learning outcome Results are presented in accordance with the structure of the interview guide presented in 2.2. The project as a whole: The two engineering groups both said that the time allocation was more than sufficient, and had it not been for printing time, this part of the assignment could have been completed in one day"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003836_0278364919835606-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003836_0278364919835606-Figure8-1.png",
        "caption": "Fig. 8. When commanding an upright resting posture with the CoM placed in the middle of the support polygon, spring compressions lead to a different posture. Ankle spring compressions make the robot lean forward whereas knee spring compressions make it lean backward. These two effects cancel each other to some extent, but, in practice, the torso would tilt depending on the desired knee angles and CoP locations. The actual CoM might also move backward for a few centimeters.",
        "texts": [
            ", 2011)), we perform position control only on the motor shafts (and not on the link positions). The term \u2018\u2018ghost posture\u2019\u2019 in Figure 1 also refers to motor shaft positions u(t) that are stiffly controlled by high proportional gains. In the following section, we discuss our joint-space control components that realize the desired Cartesian tasks on the real hardware. As mentioned previously, the SEAs of COMAN are stiff enough to keep balance, but spring compressions may produce an error in tracking. Imagine we command a crouched standing posture shown in Figure 8 with the CoM placed in the middle of the support polygon. Owing to the ankle springs, the robot tends to lean forward, whereas owing to the knee springs, the robot leans backward. These two effects slightly cancel each other, but, in practice, the knee deflection systematically perturbs our state estimation considerably. This is because the relative horizontal foot-pelvis position highly depends on the knee angle (with a radius equal to the shank length), but much less on the ankle angle. Thus, the mismatch between actual and ghost postures is considerable and needs to be considered in the control design",
            " Figure 17(A) and (B) show snapshots taken at phase change moments when forward and backward pushes are applied. Depending on the push timing, a small footstep adjustment takes place in the same phase with the push, and the next footstep locations recover the push gradually. Owing to the safety threshold applied on footstep lengths, and depending on the push strength, the recovery process might take more than two steps. Note that we typically command the pelvis in the middle of the support polygon. The ankle springs, therefore, compress slightly as shown in Figure 8 due to a non-zero ankle torque (when the CoP is not exactly under the ankle joint). Owing this default compression, the passive CoP modulation strategy resists more against forward pushes than backward, although most of the stabilization comes from the stepping strategy. Our control design and safety thresholds target small, yet dynamic footstep adjustments that do not violate the linearity of the 3LP model, sagittal\u2013lateral decoupling, and actuator limitations. When a strong push is applied, some of these assumptions might not be valid anymore, though the robot may still recover from the push over multiple steps"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001015_00423114.2017.1296168-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001015_00423114.2017.1296168-Figure3-1.png",
        "caption": "Figure 3. Wheel\u2013rail contact coordinates.",
        "texts": [
            " Considering the formal wear calculation is quite a time-consuming process, we have chosen to use wheel\u2013rail contact distribution instead as the alternative criterion to evaluate wheel\u2013rail wear rate. (4) The overall vehicle dynamic behaviour is determined by the EC vs. wheelset lateral displacement function which is derived from RRD function. Mathematical modelling and design procedure To begin with the mathematical modelling, one should first determine the global and local coordinate systems, as shown in Figure 3. Here, we use right-handed coordinate systems, as one can see from Figure 3, the abscissa of rail local coordinate system is parallel to the tangent line of rail top. Its origin is set at the centre of the track. The abscissa of wheel local coordinate system is parallel to the wheelset axis. Its origin is set at the wheelset gravity centre. Treat rail local coordinate system as the wheel\u2013rail global coordinate system, so the global coordinates of the rail origin (Or) are (0,0) and global coordinates of the wheel origin (Ow) are (yw0,zw0). When the wheelset has a lateral movement of yw and contacts with rail on both sides, it has a rolling angle \u03c6w, while the local coordinates of left and right contact points A and B in the wheel and rail local coordinate systems, respectively, are (ywl,zwl)/(yrl,zrl) and (ywr,zwr)/(yrr,zrr).Draw lineACparallel to the abscissa of thewheelset coordinate system and line AD parallel to the abscissa of the rail coordinate system. It can be deduced that \u2220CAD = \u03c6w. Normally, the wheelset rolling angle is very small (\u03c6w < 5\u00b0), thus the following relationship can be obtained through Figure 3: yrl(yw) = ywl(yw) \u2212 zwl(yw) \u00b7 \u03c6w(yw) + yw, (3) yrr(yw) = ywr(yw) \u2212 zwr(yw) \u00b7 \u03c6w(yw) + yw, (4) zwl(yw) \u2212 zwr(yw) = \u03c6w(yw) \u00b7 (ywr(yw) \u2212 ywl(yw)) + (zrl(yw) \u2212 zrr(yw)). (5) RRD function definition: R(yw) = zwl(yw) \u2212 zwr(yw). (6) Equation (5) can be written as follows: R(yw) = \u03c6w(yw) \u00b7 (ywr(yw) \u2212 ywl(yw)) + (zrl(yw) \u2212 zrr(yw)). (7) Since the wheel contact point and rail contact point is the same point in the contact space and due to the smoothness of the wheel and rail profile, there must exists a common tangent plane and a common normal line at this point that makes gradients with both profiles and the wheelset\u2019s rolling angle satisfies the following equations: dzrl(yw) dyrl(yw) = dzwl(yw) dywl(yw) + \u03bbl(yw) \u00b7 \u03c6w(yw), (8) dzrr(yw) dyrr(yw) = dzwr(yw) dywr(yw) + \u03bbr(yw) \u00b7 \u03c6w(yw), (9) where \u03bbl(yw) and \u03bbr(yw) are compensation factors varying from 1 to 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001934_1.4030344-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001934_1.4030344-Figure15-1.png",
        "caption": "Fig. 15 Magnetic bearing setup for dynamic condition",
        "texts": [
            " In the following section, comparison of dynamic performance of the best two configurations having higher load carrying capacity, i.e., full ring stator magnet and RMD configuration, is carried out. In order to establish the characteristics of the full ring PMB and proposed RMD configuration, both static and dynamic analyses of these bearing configurations are necessary. This section presents the dynamic analysis of these PMB configurations. For determining the dynamic characteristics of the bearings, a bearing setup with a disk of mass 1.5 kg was attached to shaft (shown in Fig. 15) to induce the dynamic disturbances to the system. For comparing the dynamic performance, the orbit of the rotor with respect to center of the bearing and amplitude of the vibration was measured and compared. The orbit of the rotor was measured using the displacement sensor in horizontal and vertical directions. The displacement signal provides the deviation of the rotor position from the center of the stator. The plotting of the horizontal and vertical displacement signals at a particular frequency gives the orbit plot of the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000169_ext.12016-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000169_ext.12016-Figure8-1.png",
        "caption": "Figure 8 Experimental setup at the gear rolling machine.",
        "texts": [],
        "surrounding_texts": [
            "The component material is 16 MnCr5. Tests were performed with a feed rate of 0.05 mm/s. For every 0.18\u25e6 tool rotation angle, one measurement was recorded as a result of the rotational speed of 6 rpm and a sampling rate of 200 Hz."
        ]
    },
    {
        "image_filename": "designv11_13_0000013_1.4029241-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000013_1.4029241-Figure1-1.png",
        "caption": "Fig. 1 Planar relative motion of missile and target",
        "texts": [
            " Stability of the system with the designed guidance law is analyzed. Simulation results are proposed to demonstrate the excellent property of the designed guidance law. The proposed guidance law with finite time convergence is a practical guidance law to deal with the lag of missile autopilot. The high-order derivatives of the LOS angle are avoided in the expression of guidance law such that it can be implemented in practical applications. The planar relative motion of a missile and a target is shown in Fig. 1. The missile and the target are denoted by M and T, respectively. The relative range between the missile and target is denoted by R, which is positive. The derivative of R with respect to time is denoted as _R. The velocities of the target and missile are denoted by Vt and Vm, respectively. The LOS angle is denoted by q and the derivative of q with respect to time is denoted as _q. The Contributed by the Dynamic Systems Division of ASME for publication in the JOURNAL OF DYNAMIC SYSTEMS, MEASUREMENT, AND CONTROL",
            " Journal of Dynamic Systems, Measurement, and Control MAY 2015, Vol. 137 / 051014-1 Copyright VC 2015 by ASME Downloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use flight path angles of the target and missile are denoted by ut and um, respectively. Assume that the missile and target are point masses moving in the plane and the velocities of the missile and target are constant. Then, the relative motion shown in Fig. 1 can be expressed by the following equations: _R \u00bc Vt cos\u00f0q ut\u00de Vm cos\u00f0q um\u00de (1) R _q \u00bc Vt sin\u00f0q ut\u00de \u00fe Vm sin\u00f0q um\u00de (2) Let VR \u00bc _R and Vq \u00bc R _q. Substituting them into Eqs. (1) and (2) and then differentiating them with respect to time yields _VR \u00bc V2 q R \u00fe aTR aMR (3) _Vq \u00bc VRVq R \u00fe aT aM (4) where aTR \u00bc _Vt cos\u00f0q ut\u00de \u00fe Vt _ut sin\u00f0q ut\u00de and aMR \u00bc _Vm cos \u00f0q um\u00de \u00fe Vm _um sin\u00f0q um\u00de denote the accelerations of the target and missile along the LOS, respectively; aT \u00bc Vt _ut cos\u00f0q ut\u00de _Vt sin\u00f0q ut\u00de and aM \u00bc Vm _um cos\u00f0q um\u00de _Vm sin\u00f0q um\u00de denote the accelerations of the target and missile normal to the LOS, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001715_j.isatra.2015.05.002-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001715_j.isatra.2015.05.002-Figure1-1.png",
        "caption": "Fig. 1. The schematic of the EHA.",
        "texts": [
            " The proposed observer can serve as a basis for future output feedback controllers to improve the tracking performance of position controllers for the EHA system. The rest of the paper is organized as follows: the EHA modeling issues and the problem formulation are presented in Section 2. Section 3 presents the SDDOB for the EHA system. Several numerical simulations are presented in Section 4 to show the effectiveness of the proposed observer. Section 5 contains the conclusion and the future works. The schematic of the Electro-Hydraulic Actuator studied in this paper is depicted in Fig. 1 [20,21]: The EHA system contains three parts namely the electrical, the mechanical and the hydraulic parts. These parts represents an interconnected subsystem in such a way that the dynamic of each subsystem influences the dynamics of the others. Please cite this article as: Sofiane AA. Sampled data observer based inter-sample output predictor for Electro-Hydraulic Actuators. ISA Transactions (2015), http://dx.doi.org/10.1016/j.isatra.2015.05.002i The electrical part of the EHA system is a servo-valve which controls the entire EHA system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002895_s1061920816040026-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002895_s1061920816040026-Figure3-1.png",
        "caption": "Fig. 3. Euler angles.",
        "texts": [],
        "surrounding_texts": [
            "A change of fixed actions OX1X2X3 \u2192 OX \u2032 1X \u2032 2X \u2032 3 is also defined by an orthogonal matrix\nR = SR\u2032, S \u2208 SO(3); as a result, relation (3) is represented in the form R\u2032 = Q\u2032r, Q\u2032 = SQ.\nA change of moving axes Ox1x2x3 \u2192 Ox\u2032 1x \u2032 2x \u2032 3, defined by r\u2032 = Sr, S \u2208 SO(3), gives, corre-\nspondingly, R = Q\u2032r\u2032, Q\u2032 = QS.\nFrom the geometric viewpoint, a change of fixed axes defines a left translation on the group SO(3), while a change of moving axes defines a right translation on the group.\n2.2. Left-Invariant Metric on the Group and the Kinetic Energy of the Body\nIf the orientation of the body depends on time, Q(t), then, by (3), for every point r = (x1, x2, x3) of the body B, the velocity of this point relative to the fixed axes is given by the relation\nR\u0307 = Q\u0307r. (4)\nAs is well known, the pair (Q, Q\u0307) defines a point in the tangent bundle TSO(3) and completely characterizes the state of the body (i.e., the orientation and velocity). An important role in the derivation of the equations of motion of the body is played by a special function on TSO(3), namely,\nthe kinetic energy T (Q, Q\u0307) (which is the sum of the kinetic energies of the elements forming the body). Using the relation (4), we represent T in the form\nT =\n\u222b\nr\u2208B\n( R\u0307(r), R\u0307(r) ) \u03c1(r)d3r = 1\n2 Tr(Q\u0307JQ\u0307T ). (5)\nHere J is the 3\u00d7 3 symmetric matrix with the entries\nJij =\n\u222b\nr\u2208B \u03c1(r)xixjd\n3r,\nwhere \u03c1(r) stands for the density of the body at the point r.\nThe left and right actions on the group SO(3) can naturally be extended to TSO(3),\n\u2013 left action: LS(Q, Q\u0307) = (SQ,SQ\u0307), S \u2208 SO(3). \u2013 right action: RS(Q, Q\u0307) = (QS, Q\u0307S),\nObviously, the kinetic energy does not depend on the choice of the fixed axes, and therefore, from the point of view of the action of the group on itself, we obtain the following assertion.\nProposition 1. If the matrix of the direction cosines is defined as in (2), then the kinetic energy of the rigid body is invariant under the left translations on SO(3).\nProof. Let a left translation on SO(3) be defined by the relation\nQ\u2032 = SQ;\nthen, using the properties of the trace, we obtain\nT \u2032 = 1\n2 Tr(SQ\u0307JQ\u0307TST ) =\n1 2 Tr(Q\u0307JQ\u0307T ) = T.\nFrom the geometric point of view, the kinetic energy of a rigid body with the fixed point (5) defines a left-invariant metric on the group SO(3).\nRemark. In general, the kinetic energy is not invariant under the right translations Q = QS:\nT \u2032 = 1\n2 Tr(Q\u0307J\u2032Q\u0307T ), J\u2032 = SJST .\nRUSSIAN JOURNAL OF MATHEMATICAL PHYSICS Vol. 23 No. 4 2016",
            "2.3. Parametrization, Quasivelocities, and Left-Invariant Vector Fields on the Group\nLet us parametrize the group SO(3) by local coordinates q = (q1, q2, q3) for which we take the Euler angles q1 = \u03b8, q2 = \u03d5, q3 = \u03c8. In this case, a general rotation of the body Q \u2208 SO(3) is realized as a sequence of three rotations,\nQ1 = ( cos\u03c8 \u2212 sin\u03c8 0 sin\u03c8 cos\u03c8 0 0 0 1 ) , Q2 = ( 1 0 0 0 cos \u03b8 \u2212 sin \u03b8 0 sin \u03b8 cos \u03b8 ) , Q3 = ( cos\u03d5 \u2212 sin\u03d5 0 sin\u03d5 cos\u03d5 0 0 0 1 ) ,\nQ = Q1Q2Q3\n= ( cos\u03d5 cos\u03c8\u2212cos \u03b8 sin\u03d5 sin\u03c8 \u2212 sin\u03d5 cos\u03c8 \u2212 cos \u03b8 cos\u03d5 sin\u03c8 sin \u03b8 sin\u03c8 cos\u03d5 sin\u03c8+cos \u03b8 sin\u03d5 cos\u03c8 \u2212 sin\u03d5 sin\u03c8 + cos \u03b8 cos\u03d5 cos\u03c8 \u2212 sin \u03b8 cos\u03c8\nsin \u03b8 sin\u03d5 sin \u03b8 cos\u03d5 cos \u03b8\n) .\nThe first rotation by the angle \u03c8 (the angle of precession) around the axis OX3 takes the moving trihedron Ox1x2x3 to the position Ox1 \u2032x2 \u2032x3\n\u2032. The second rotation by the angle \u03b8 (the angle of nutation) is carried out around the axis Ox1\n\u2032, the so-called nodal line. The last rotation by the angle \u03d5 (the angle of intrinsic rotation) around the Ox3 axis brings the two trihedra into coincidence.\nAs we have seen above, the kinetic energy of the body (5) is a left invariant function on TSO(3),\nwhile the (generalized) velocities q\u0307 = (\u03b8\u0307, \u03d5\u0307, \u03c8\u0307) are not invariant under left (and right) translations on the group. Therefore, analyzing rigid body dynamics, as a rule, it is more convenient to use leftinvariant quasivelocities, for which one takes the components of the angular velocity \u03c9 = (\u03c91, \u03c92, \u03c93) in the moving coordinate system Ox1x2x3. In the matrix form, these components are defined in the standard way,\n\u03c9\u0302 = Q\u22121Q\u0307 = ( 0 \u2212\u03c93 \u03c92 \u03c93 0 \u2212\u03c91\n\u2212\u03c92 \u03c91 0\n) =\n=\n\u239b \u239d 0 \u2212\u03d5\u0307\u2212 \u03c8\u0307 cos \u03b8 \u2212\u03b8\u0307 sin\u03d5+ \u03c8\u0307 sin \u03b8 cos\u03d5\n\u03d5\u0307+ \u03c8\u0307 cos \u03b8 0 \u2212\u03b8\u0307 cos\u03d5\u2212 \u03c8\u0307 sin \u03b8 sin\u03d5 \u03b8\u0307 sin\u03d5\u2212 \u03c8\u0307 sin \u03b8 cos\u03d5 \u03b8\u0307 cos\u03d5+ \u03c8\u0307 sin \u03b8 sin\u03d5 0\n\u239e \u23a0 .\n(6)\nThis enables one to express the kinetic energy (5) using the angular velocity of the body,\nT = \u22121\n2 Tr(\u03c9\u0302J\u03c9\u0302).\nIn mechanics, instead of the tensor J, it is customary to use the inertia tensor of the form\nI = (TrJ)E \u2212 J;\nhere, in the vector form, the kinetic energy is represented as\nT = 1\n2 (\u03c9, I\u03c9).\nRUSSIAN JOURNAL OF MATHEMATICAL PHYSICS Vol. 23 No. 4 2016",
            "Let us express the generalized velocities using the relations (6),\n\u03b8\u0307 = \u03c91 cos\u03d5\u2212 \u03c92 sin\u03d5, \u03c8\u0307 = \u03c91 sin\u03d5\nsin \u03b8 + \u03c92\ncos\u03d5 sin \u03b8 ,\n\u03d5\u0307 = \u2212\u03c91 cos \u03b8 sin\u03d5\nsin \u03b8 \u2212 \u03c92\ncos \u03b8 cos\u03d5\nsin \u03b8 + \u03c93.\n(7)\nUsing this, we see that the evolution of every function f(q) on G = SO(3) can be represented in the form\ndf(q)\ndt = 3\u2211 i=1 q\u0307i \u2202f \u2202qi = 3\u2211 \u03b1=1 \u03c9\u03b1\u03c4\u03b1(f),\n\u03c41 = cos\u03d5 \u2202 \u2202\u03b8 \u2212 cos \u03b8 sin\u03d5 sin \u03b8 \u2202 \u2202\u03d5 + sin\u03d5 sin \u03b8 \u2202 \u2202\u03c8 ,\n\u03c42 = \u2212 sin\u03d5 \u2202 \u2202\u03b8 \u2212 cos \u03b8 cos\u03d5 sin \u03b8 \u2202 \u2202\u03d5 + cos\u03d5 sin \u03b8 \u2202 \u2202\u03c8 , \u03c43 = \u2202 \u2202\u03d5 .\n(8)\nHere \u03c4\u03b1, \u03b1=1,2,3 are left-invariant vector fields on SO(3) forming the Lie algebra so(3) of SO(3), [\u03c41, \u03c42] = \u03c43, [\u03c42, \u03c43] = \u03c41, [\u03c43, \u03c41] = \u03c42.\nFrom the geometric point of view, the motion of a rigid body is defined by a curve q(t) on the group G = SO(3), where the coordinates of the angular velocity \u03c9 = (\u03c91, \u03c92, \u03c93) in the moving axes Ox1x2x3 are the components of the velocity of the curve q(t) in the basis of the Lie algebra g of left-invariant vector fields,\n\u03c4\u03b1 = \u2211 i \u03c4\u03b1i \u2202 \u2202qi , [\u03c4\u03b1, \u03c4\u03b2 ] = \u2211 \u03b3 c\u03b3\u03b1\u03b2\u03c4\u03b3 ,\nwhere c\u03b3\u03b1\u03b2 are the structure constants of the Lie algebra and the coefficients \u03c4\u03b1i are defined by the transformation\nq\u0307i = \u2211 \u03b1 \u03c4\u03b1i\u03c9\u03b1.\nSimilarly, it turns out that the projections of the angular velocity to the fixed axes \u03a9 = (\u03a91,\u03a92,\u03a93) define the quasivelocities invariant under the right translations (i.e., changes of the moving axes). Their matrix representation is of the form\n\u03a9\u0302 = Q\u0307Q\u22121 = ( 0 \u2212\u03a93 \u03a92 \u03a93 0 \u2212\u03a91\n\u2212\u03a92 \u03a91 0\n)\n=\n\u239b \u239d 0 \u2212\u03d5\u0307 cos \u03b8 \u2212 \u03c8\u0307 \u03b8\u0307 sin\u03c8 \u2212 \u03d5\u0307 sin \u03b8 cos\u03c8\n\u03d5\u0307 cos \u03b8 + \u03c8\u0307 0 \u2212\u03b8\u0307 cos\u03c8 \u2212 \u03d5\u0307 sin \u03b8 sin\u03c8 \u2212\u03b8\u0307 sin\u03c8 + \u03d5\u0307 sin \u03b8 cos\u03c8 \u03b8\u0307 cos\u03c8 + \u03d5\u0307 sin \u03b8 sin\u03c8 0\n\u239e \u23a0 .\nThe corresponding right-invariant vector fields\nT1 = cos\u03c8 \u2202\n\u2202\u03b8 +\nsin\u03c8\nsin \u03b8\n\u2202 \u2202\u03d5 \u2212 cos \u03b8 sin\u03c8 sin \u03b8 \u2202 \u2202\u03c8 ,\nT2 = sin\u03c8 \u2202 \u2202\u03b8 \u2212 cos\u03c8 sin \u03b8 \u2202 \u2202\u03d5 + cos \u03b8 cos\u03c8 sin \u03b8 \u2202 \u2202\u03c8 , T3 = \u2202 \u2202\u03c8\ndefine the commutators of the form [T1,T2] = \u2212T3, [T2,T3] = \u2212T1, [T3,T1] = \u2212T2. (9)\nWe point out the well-known natural reciprocity property of left and right-invariant vector fields on the group.\nProposition. Let \u03c4\u03b1 and T\u03b1, \u03b1 = 1, . . . , n, be, respectively, left-invariant and right-invariant fields on the group. Then, for arbitrary \u03b1, \u03b2 = 1, . . . , n, we have\n[\u03c4\u03b1,T\u03b2 ] = 0. (10)\nRUSSIAN JOURNAL OF MATHEMATICAL PHYSICS Vol. 23 No. 4 2016"
        ]
    },
    {
        "image_filename": "designv11_13_0001585_aim.2018.8452392-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001585_aim.2018.8452392-Figure8-1.png",
        "caption": "Figure 8. Peristaltic crawling robot \u201cPEW-RO V\u201d",
        "texts": [
            " In addition, solenoid valves are stored inside the joint. The robot was also problematic in that the amount of contraction in the unit part alone was small, and the air pressure response was poor. The robot was difficult to handle when put to practical use. A peristaltic motion-type robot was developed with the aim of improving the motion of the robot with a view toward practical applications and improving basic performance. The authors developed another peristaltic motion robot named PEW-RO V. A schematic of PEW-RO V is shown in Fig.8. The specifications of the robot are shown in Table 3. Fig.10 shows the contraction of the unit part in the pipe. The amount of contraction inside the pipe depends only on the diameter of the artificial muscle and the diameter inside the pipe (Eq. (1)). Compared with the previous unit part, the diameter of the artificial muscle is thinned by reducing the size of the artificial muscle fastener (Fig.11). As a result, it was confirmed that the contraction inside the tube improved by 70%. The contraction characteristics of the existing robot unit are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000647_iciea.2015.7334244-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000647_iciea.2015.7334244-Figure4-1.png",
        "caption": "Fig. 4 Collision detection of sphere-cylinder",
        "texts": [],
        "surrounding_texts": [
            "Keywords-collision detection; manipulator; vector relation\nI. INTRODUCTION Collision detection plays an important role in manipulator field, which includes manipulator-environment collision and manipulator-manipulator collision. It is necessary to build proper geometric models to detect collision of manipulator rapidly. Therefore, how to model and how to process model is directly related to the precision and efficiency of collision detection algorithm.\nConsidering the complexity of manipulator\u2019s shape and the diversity of environment, and in order to reduce mathematical operator to ensure efficiency, bounding box algorithm is generally used in collision detection. Basic idea of the algorithm is to use a slightly larger geometry with simple feature(called the bounding box) to represent complex objects approximately. Simple geometry such as sphere and cuboid are widely used. Angel P. del Pobil [1] used hierarchies of detail based on spheres to represent complex objects, which was of high efficiency. Fan Ouyang [2] analyzed the advantages of hierarchical structure and octree data structure, and proposed an octree-based spherical hierarchical model for collision detection. Crnekovic Mladen [3] took the environment as a series of sphere or prism and compared the efficiency and error between the two models. Kei Okada [4] adopted AABB algorithm to envelope objects as closely as possible, and demonstrated real-time collision detection. Zonggao Mu [5] turned to OBB algorithm considering that most space manipulators are composed of modular and standardized\ndevices, and developed an integrated simulation system based on OSG [6,7] to verify the algorithm. Byungjin Jung [8] used a kind of band pass filter to avoid false alarm in collision detection, and the experimental result showed that the method is not only fast but also robust. Apart from bounding box algorithm, Ruining Huang [9] adopted a hierarchical bounding box method and space decomposition method comprehensively. Minxiu Kong [10] proposed a collision detection algorithm for coordinated industrial robots based on calculating the shortest distance between two geometry.\nObviously, these three bounding boxes have their own strong points and weakness. Among them, Sphere box is the simplest with high efficiency for collision detection, although with low precision at the same time. AABB is acceptable in both efficiency and precision. OBB is of high precision, but it need to detect 15 separate axes to determine collision status in 3D environment. To ensure both precision and efficiency, in this paper a rapid collision detection algorithm for manipulator based on the vector relation of point, line segment and rectangle is proposed. Besides, we put forward the concept of \u201ccenter distance\u201d and \u201cgeometric dimension\u201d, and provided a center distance criterion to detect collision rapidly.\nII. GEOMETRY OF MANIPULATOR AND ENVIRONMENT First, we simplified manipulator and environment based on the bounding box theory. Considering that most manipulators consist of modular components, we adopted cylinder and cuboid to represent a manipulator. For diverse obstacles in environment, we used sphere, cylinder, cuboid and combination of them to represent.\n934978-1-4799-8389-6/15/$31.00 c\u00a92015 IEEE",
            "Therefore, object collision is simplified to geometry collision. In this paper, geometry is subdivided as the combination of point, line segment (denoted as line for convenience) and rectangle. Six kinds of geometry (point, line, rectangle, sphere, cylinder and cuboid) above are collectively called basic geometry, as shown in Fig. 2.\n0P\ncy\ncx\ncz\ncO\n2P\n1P 3P 4P\n5P\n6P 1s\n2s3s\n1P\n2P cO 1s\n2s\n0P\n1P\n2P r\nr\nFig. 2 Basic geometry\nFurthermore, we extracted concept of \u201ccenter distance\u201d \u2013 the distance between center of two geometry; and \u201cgeometry dimension\u201d \u2013 the maximum distance from center to surface of a geometry. The information of all basic geometry is shown in Table 1.\nIII. COLLISION DETECTION OF BASIC GEOMETRY After simplification of manipulator and environment, object collision can be classified into 4 kinds of collision \u2013 spherecylinder, sphere-cuboid, cylinder-cylinder and cuboid-cuboid collision. For convenience, cylinder-cuboid collision is seen as cuboid-cuboid collision. To skip collision detection between 2 geometry far away, we proposed a center distance criterion: if the center distance between 2 geometry is greater than the sum of their geometric dimension, there is no collision. Otherwise it need further detection. In addition, considering safety margin dS , the criterion is shown with following expression:\n( , ) . . i j i j Sif Dist G G D D d No collision else Need more detection > + +\nIn further detection, the 4 types of collision detection can be converted into 5 basic detections: point-line, point-cuboid, lineline, line-rectangle, line-cuboid. Their relation is given in Fig. 3.\nThe following is an introduction of the 4 types of collision detection. The algorithm returns 1 if any collision happens,; otherwise returns 0.\na) Apply center distance criterion on sphere-cylinder. The algorithm returns 0 if (1) meets, otherwise turns to b);\n2 2 0 1 2( ) = || || || || /4S C m C S SDist G ,G P P PP r r d> + + + (1)\nb) Compute the shortest distance dmin between point P0 to line P1P2.\nDo projection of point P0 to line P1P2. If 1 1 2/ [0,1]vPP PP\u03bb = \u2208 , then min 0|| ||vd P P= , otherwise\nmin 0 1 0 2 min{ || ||,|| ||}d P P P P= . c) Detect collsion of sphere-cylinder by dmin. The algorithm returns 0 if (2) meets; otherwise returns 1.\na) Apply center distance criterion on sphere-cuboid. The algorithm returns 0 if (3) meets, otherwise turns to b);\n2 2 2 0 1 2 3Dist( , ) || || || || || || || ||S C c S SG G P O r d= > + + + +s s s (3)\nb) Detect whether point P0 is outside of cuboid C. For any surface of cuboid with outward normal vector in\nand center point iP , if 0 0i iP P n\u22c5 < , then point P0 is outside, otherwise turns to c). The algorithm computes dmin of point P0 to the corresponding surface. The algorithm returns 0 if dmin > rS ; otherwise returns 1.\nc) Detect next surface.\n2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA) 935",
            "If all the 6 surfaces of cuboid fail to meet 0 0i iP P n\u22c5 < , then point P0 is inside of cuboid, which means 1 to return. 3) Cylinder \u2013 Cylinder\n1P\n2P\n1vP\nmind\n3P\n4P\n2vP\n1C\n2C\n2l 1l\n1mP 2mP\n1Cr 2Cr\n1mP\n4P\n2mP\n1vP\n2vP\nmind\n2P\n1P\n1 3 1 4 2 3 2 4 min{|| ||,|| ||,|| ||,|| ||}PP PP P P P P . c) Detect collsion of cylinder-cylinder by dmin. The algorithm returns 0 if (7) meets; otherwise returns 1.\na) Apply center distance criterion on cuboid-cuboid: The algorithm returns 0 if (8) meets, otherwise turns to b);\n2 2 2 1 2 1 2 3\n2 2 2 1 2 3\nDist( , ) || || || || || || || ||\n|| || || || || ||\nC C c c\nS\nG G O O '\n' ' ' d\n= > + +\n+ + + +\ns s s\ns s s (8)\nb) Detect whether vertex of C1 is inside of C2. The collision is the point-cuboid type. The algorithm\nreturns 1 if a vertex collides, otherwise detects next vertex. If all 8 vertexes have no collision, turns to c);\nc) Detect whether edge of C1 is in collision with C2. The collision is the line-cuboid type. The algorithm\nreturns 1 if an edge collides, otherwise detects next edge. If all 12 edges have no collision, returns 0.\nIV. COLLSION DETECION OF MANIPULATOR In the base of simplified geometric models and collision detection of basic geometry, we realized collision detection of manipulator. The algorithm flow chart is shown in Fig. 8.\nThe simplified model of a n-DOF manipulator is shown in Fig. 9.\nTo process collision detection of manipulator, we need the key information of basic geometry \u2013 key point, key vector and key length in an absolute reference system. In this paper, we took the base coordinate system of manipulator as the absolute reference system. Different geometry demands for different key information, as listed in Table 2. The expression of key point and key vector are same in Cartesian form. When joint angles are given, the relevant configuration, key point and key vector of manipulator can be solved by forward kinematics. The key length will not change without considering flexibility of manipulator.\n936 2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA)"
        ]
    },
    {
        "image_filename": "designv11_13_0000369_iros.2013.6696872-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000369_iros.2013.6696872-Figure5-1.png",
        "caption": "Fig. 5. Two bipedal models with torso. There are two hip actuators, each acting between the torso and the corresponding thigh. The swing retraction torque, quantified by the impulse R, decreases \u03c6 and pushes the swing leg toward the stance leg. (a) The straight-leg model: the push-off impulse P is provided by a prismatic actuator along the stance leg. (b) The articulatedleg model: push-off is provided by the knee and ankle torques, quantified by their impulses K and A, which tend to extend the corresponding angles (consequently extending the leg).",
        "texts": [
            " The resulting optimal relative timing is achieved based on the fact that J\u03c9/P =\u2212Jv/R< 0. The equality relation here is guaranteed by the symmetry of the mass matrix, and the inequality comes from the fact that the push-off force tends to retract the swing leg. Interestingly, these conditions are satisfied for even more complex models, implying that the same optimal relative timing of impulsive push-off and retraction holds for a large range of bipedal models. For example, consider the more realistic models in Fig. 5. In these models the retraction impulse R is applied by the swing hip actuator (both models have two hip actuators acting between the torso and the corresponding leg) and pushes the swing leg toward the stance leg (decreasing the swing hip angle \u03c6). In the model with straight legs the push-off impulse P is directly applied by the prismatic actuator along the stance leg, whereas in the model with articulated legs the equivalent push-off impulse is provided by the knee impulse K and the ankle impulse A that extend the corresponding joint angles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003033_tcst.2017.2695161-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003033_tcst.2017.2695161-Figure1-1.png",
        "caption": "Fig. 1. Schematic of the transect-following controller.",
        "texts": [
            " Let T = [T1, T2]T be a unit 2-D vector that defines the direction of the travel for the transect. Define a matrix J = [ 0 1 \u22121 0 ] and let N be the unit vector N = JT. Then, N is obtained by rotation T counterclockwise by \u03c0/2, and hence, N is orthogonal to T. In the coordinate frame formed by T and N, the velocity of the vehicle v can be represented as v(x, t) = (v(x, t)T N)N + (v(x, t)T T)T and the modeled flow velocity can be represented as FM (x, t) = (FM (x, t)T N)N + (FM (x, t)T T)T. Let p \u2208 be an arbitrary point on the transect, as shown in Fig. 1. The control effort u determines the heading of the vehicle toward direction [cos u, sin u]T , so that v = [s cos u, s sin u]T , where s is the through-water speed. The control u will be designed as a feedback law to cancel modeled flow velocity orthogonal to the transect direction, and an additional proportional controller is used to maintain the vehicle on the transect line. Let satc(\u00b7) denote the saturation function, satc : R \u2192 [\u2212c, c], where satc(y) = \u23a7\u23aa\u23a8 \u23aa\u23a9 \u2212c, if y < \u2212c y, if \u2212 c \u2264 y \u2264 c c, if y > c",
            " (12) Using the notation V(x, t; u) = FM (x, t) + v(u(x, t)) defined in (6), we have V(x, t; u) = (FT M T+\u221a\u03b3 (x) ) T\u2212K ((x\u2212p)T N)N (13) if \u03b3 \u2265 0 and V(x, t; u) = ( FT M T ) T + (FT M N \u2212 sign ( FT M N + K (x \u2212 p)T N ) s ) N (14) otherwise. When the vehicle speed s is larger than the cross-track component of the flow, then the vehicle cancels the cross-track flow and applies proportional-gain control to return to the transect. If the vehicle speed is weaker than the cross-track component of the flow, then the vehicle aims along the direction \u2212N shown in Fig. 1 toward the transect. Like the transect-following controller, the station-keeping controller is defined for a constant-speed vehicle model with heading control u. The vehicle should always move toward a fixed goal position g. Thus, the heading u is chosen to cancel the modeled flow velocity in the direction normal to the desired motion. Let T = [T1, T2]T be the unit vector from the vehicle position to the goal g. Without loss of generality, we choose the coordinate system so that g = 0. Then, T(x) = \u2212(x/\u2016 x \u2016)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003373_s0005117918030062-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003373_s0005117918030062-Figure1-1.png",
        "caption": "Fig. 1. (a) Coordinate system related to the body is shown on the figure by axes x,y, z. Propellers 1 and 3 rotate in the direction opposite to the rotation of propellers 2 and 4. The air resistance moment (braking moment) \u03c4d acts in the direction opposite to the angular velocity of the vehicle \u03c9B BE expressed in the coordinate system related to the body \u03c9B BE =\u0307 (p, q, r). The vehicle is subject to the gravity force mg. (b) We separately show the ith propeller which is rotating with speed \u03c9i with respect to the body. Each propeller creates the thrust fi and moment \u03c4i, both in the direction of the propeller\u2019s rotation axis. In this case \u03c9i > 0, \u03c4i < 0.",
        "texts": [],
        "surrounding_texts": [
            "As part of the general problem of constructing an emergency landing algorithm in case of type 1 and 4 failures for a quadcopter, we have to introduce assumptions for constructing an adequate mathematical model of a quadcopter. We divide them in two groups. The first group includes assumptions caused by constraints that arise from the level of technology development. For example, at the modern level of technology we cannot create a quadcopter of mass 1 kg and duration of flight 5 h since a 1 kg battery simply would not be able to store enough energy for 5 h of flight. Thus, we have the following assumptions. \u2022 Maximal thrust does not exceed twice the force of gravity for the vehicle, and minimal thrust during flight does not fall below the value defined at idle Um > 0. Thus Um < fi 2mg = UM , i = 1, 3. One usually considers the model of a quadcopter that has an upper bound on the control 1.5 times larger, but there are very few real models with these characteristics. \u2022 The vehicle has weak windage, i.e., it cannot hover with propellers switched off. As a corollary, we assume that the wind does not significantly influence the flight. \u2022 The propeller reaches desired speed in some time; in the simplest case \u03c9\u0307i = 1 T\u03c9 (\u03c9U \u2212 \u03c9i), where \u03c9U is the desired angular velocity of the ith propeller, \u03c9i is the current angular velocity of the ith propeller, T\u03c9 \u2248 0.01\u2212 0.02 is the approximate time of delay. It is also important to AUTOMATION AND REMOTE CONTROL Vol. 79 No. 3 2018 distinguish the notions of mechanical and magnetic moments of the motor. The magnetic moment is defined by the number and quality of magnets on the rotor, section of the cable coiled at the stator\u2019s teeth, number of teeth, magnetic properties of the material that the teeth plates are made of, and so on. The mechanical moment, on the other hand, is defined by physical size of the motor since the propeller is attached directly to the motor\u2019s rotor, there are no losses in transition links, and due to rigid attachment and perfect balance of the propeller we can assume that there are no vibrations, however small, on the rotor. \u2022 Frequency of orientation control is 650 Hz since inertial sensors give readings with frequency 1200 Hz. Position control is done with a much lower frequency; it is only 50 Hz since the geodesic position and velocity of the vehicle are measured with this rate. Frequency data will be used for modeling the motion. Let us now consider the second group of assumptions that let us significantly simplify the original mathematical model that defines quadcopter dynamics. \u2022 The vehicle at the moment of failure is in the hovering state, i.e., there is neither explicit rotation around the vertical axis nor significant inclination of the vehicle\u2019s plane. \u2022 The reactive moment of propeller \u03c4i is linearly proportional to the thrust force, i.e., \u03c4i = (\u22121)i+1k\u03c4fi. \u2022 The thrust force of each propeller is proportional to the square of the propeller\u2019s angular velocity, i.e., fi = kf\u03c9 2 i . \u2022 The aerodynamic resistance acts only against the angular velocity of rotation with respect to Oz, i.e., \u03c4d = (\u03c4dx , \u03c4dy , \u03c4dz) = (0, 0,\u2212\u03b3zr). The quadratic term is negligible. \u2022 The control rate of propeller rotation is sufficiently high that the motor\u2019s velocity does not influence the motion of the quadcopter. Thus, we can assume that IP IB, and therefore IP \u03c9\u0307Pi 0, where \u03c9Pi = (0, 0, \u03c9i) is the angular velocity of the propeller, IP = diag(IPxx, I P yy, I P zz) and IB = diag(IBxx, I B yy, I B zz) are matrices of inertia for the propellers and quadcopter respec- tively. \u2022 The moment of rotation created by the motor\u2019s rotor can be comparable with the moment of rotation of the body, so the value IP\u03c9Pi cannot be treated as negligible despite the previous item. \u2022 The body of a quadcopter has symmetry, i.e., IBxx = IByy. Here the dynamical model of the quadcopter has symmetric propeller locations with respect to the center of mass."
        ]
    },
    {
        "image_filename": "designv11_13_0002081_j.electacta.2014.12.148-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002081_j.electacta.2014.12.148-Figure10-1.png",
        "caption": "Fig. 10: SEM images of (A) PLG/GC and (B) poly/PLG/PLG/GC after 25 cycles.",
        "texts": [
            " ((place Figure 7 here)) ((place Figure 8 here)) Independent of the electrode nature and the layer thickness, the surface structure of polymers appeared to be homogeneous and smooth with no significant agglomeration. Figure 9A shows an example of a polyPLG/PLG/GC formed in 25 CV cycles. The surface topography was dominated by polishing streaks on the GC support that are smaller than 20 nm in depth. The substrate features were flattened out and lose contrast as the film thickness grows during further cycling (Fig. 9B). ((place Figure 9 here)) Complementary SEM measurements revealed a homogeneous polyPLG/PLG/GC film (Fig. 10A). Most importantly, the morphology of the polymer is charaterized by a globular morphology and pits probably affected by the grafted PLG layer (Fig. 10B). This characteristic was not observed for polyPLG/GC where a homogeneous thin layer follows more or less the topography of the electrode (Supporting Information, Fig. SI-6). ((place Figure 10 here)) Conclusion The grafting of plumbagin allows subsequent electropolymerization of a plumbagin polymer. The behavior observed during electropolymerization was different than in direct electropolymerization on GC. In contrast to polyPLG/GC, the growth of polyPLG/PLG/GC is not limited. This allows a free adjustment of the film thickness that contains active quinone groups accessible to the electrolyte. The growth of this polymer on a grafted plumbagin monolayer apparently led to the permeability of this polymer",
            " 7: (A) Example of AFM thickness determination after mechanical removal of the polymer (7 Hz and at a set-point of 0 V); (B) the corresponding cross section profile of the area marked by the white rectangle in (A). ACCEPTED M ANUSCRIP T Fig. 8: (A) AFM thickness determination of (1) polyPLG/PLG/GC after 10, 25, 50, 100, 150 potential cycles and (2) polyPLG/GC (after 5, 25 and 50 potential cycles). (B) Accumulated charge of phenoxy radical peak as function of potential cycle number for (1) polyPLG/PLG/GC and (2) polyPLG/GC. ACCEPTED M ANUSCRIP T Fig. 9: AFM topography images of the polyPLG/PLG/GC after (A) 25 and (B) 100 CV cycles. ACCEPTED M ANUSCRIP T Fig. 10: SEM images of (A) PLG/GC and (B) poly/PLG/PLG/GC after 25 cycles. ACCEPTED M ANUSCRIP T"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002692_s00500-016-2217-8-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002692_s00500-016-2217-8-Figure3-1.png",
        "caption": "Fig. 3 Bearing faults: a outer race, b inner race, c excessive wear, and d short circuit",
        "texts": [
            " In order to obtain current signals from a motor operating with defective bearings, new pairs of bearingswere subjected to the conditions that rendered them defective. Four types of faults were recreated in a controlled manner, such as an outer race fault (Zarei et al. 2014; Prieto et al. 2013; Vakharia et al. 2015; Pandya et al. 2014; Pal\u00e1cios et al. 2015), inner race fault (Zarei et al. 2014; Prieto et al. 2013; Vakharia et al. 2015; Pandya et al. 2014), short circuit and excessive wear (Prieto et al. 2013). These defects are inserted in a controlled manner, as shown in Fig. 3, with the variation of the load torque and the application of voltage unbalances in the machine feed, which are used to reproduce the operating conditions encountered in an industrial setting. This work also presents an alternative sensing method, based on the traditional use of two current sensors. Additional tests are conducted with different neural architectures in order to optimize the standards classifier structure and consequent reduction of computing costs. The use of signals in the time domain simplifies the data processing and eliminates the need for using signal analysis techniques in the frequency domain, as well as in FFT and Wavelet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003697_j.ijthermalsci.2018.11.027-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003697_j.ijthermalsci.2018.11.027-Figure1-1.png",
        "caption": "Fig. 1. Bearing temperature field distribution at rotation speed 10000r/min.",
        "texts": [
            " The contact area of the raceway and the ball were regarded as the loading surface of bearing heat flux density in the model, and the thermal convection coefficient calculated from the empirical formula are loaded as initial conditions. During the analysis, bearing speed is at 10000 r/min, and the ambient temperature in steady state is set equal to the temperature of the air around bearing housing (27 \u00b0C). The boundary condition in the finite element model is shown in Table 2. Due to the existence of rotation and spin motions, the axis and the contact area of the ball are constantly changing during the rotation. Therefore, the temperature field of the ball surface ultimately reaches an equilibrium state as shown in Fig. 1. It shows that the maximum temperature of the bearing is about 40.95 \u00b0C, located in the contact area between the inner ring and the ball. The ball surface temperature at the contact area, about 39.71 \u00b0C, is a bit lower than that of the inner raceway. The outer raceway temperature is even lower, which is 37.82 \u00b0C. As mentioned at the beginning, this paper first presented a reverse solution method for the internal thermal transfer analysis of the rolling bearing. Based on the above calculation, bearing internal convectional heat transfer parameters were obtained as the initial value"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002968_s00170-017-0058-y-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002968_s00170-017-0058-y-Figure12-1.png",
        "caption": "Fig. 12 The geometric model with BD",
        "texts": [
            " The average deformation value of two sides is obtained accurately for residual angular distortion. The results are shown in Table 3 The welding conditions are different between samples with and without BD because BD leads to increase of cross-sectional area in groove. Thus, welding process with BD must be simulated separately. The angular distortion of model without BD is calculated to be 1.9\u00b0 (Table 3). Accordingly, the reserved angle is set to be 1.5, 2.0, and 2.5\u00b0 when designing model with BD. The result obtained from the geometry modeling with a scale of 1:1 is shown in Fig. 12. Aluminum alloy sheet is placed at the bottom of the Invar alloy plate to obtain BD. Besides, the groove, root face, and space between two half plate of model with BD are the same as the model without BD. The models with different reserved angle of 1.5, 2.0, and 2.5\u00b0 were divided into hexahedral grid with a transition ratio of 1:3. The mesh model with BD is shown in Fig. 13. Models with different BD were simulated with coupled thermo-mechanical method using Marc. It is essential to mention that the material parameters, initial condition, and boundary condition were the same as the model without BD"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003549_j.jmatprotec.2018.08.017-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003549_j.jmatprotec.2018.08.017-Figure10-1.png",
        "caption": "Fig. 10. The calculated temperature fields on the top surface: (a) Experiment 1, (b) Experiment 2, (c) Experiment 3, and (d) Experiment 4 in Table 2.",
        "texts": [
            " Experimental tests were carried out to verify the accuracy of the numerical simulations. Fig. 9 shows a comparison between the experimentally obtained weld geometries and the calculated ones for the cases detailed in Table 2. Table 3 lists the measured and predicted weld dimensions for all four welding conditions. The experimental fusion zone geometry agrees reasonably well with the corresponding calculated results for all cases, indicating that the established thermal property parameters and the developed heat source model can be used to simulate the VPPA\u2013GMAW process. Fig. 10(a)\u2013(d) shows the predicted temperature distribution on the top surface for all four welding conditions. The weld width of VPPA\u2013GMAW for different welding parameters was much narrower than that of GMAW. As the VPPA power increased, the VPPA weld pool constantly expanded and gradually approached the GMAW weld pool. When the VPPA welding current was 150 A and the GMAW welding current was 100 A, the weld pools of VPPA and GMAW interacted with each other; consequently, the best coupling effect and a deep-penetration weld were achieved. Fig. 11 shows images of the weld pools for different welding conditions. The experimental weld pool shapes are consistent with the calculated weld pool shapes in Fig. 10, which further verifies the accuracy of the numerical results. Based on the predicted temperature distribution for the different welding conditions during VPPA-GMAW, it is found that the VPPA power accounting for 63% of the total power (Test Case 4) can achieve the best hybrid welding effect and full penetration. This optimized welding condition was used to achieve butt-welded joints on Al\u2013Cu\u2013Mg alloy plate with a thickness of 6mm. The morphology of the weld seam is shown in Fig. 12. There are no undercuts or hump defects on the weld seam"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002825_s40799-016-0158-x-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002825_s40799-016-0158-x-Figure1-1.png",
        "caption": "Fig. 1 (a) Schematic diagram of the MFS with the PSS and accelerometers installed in the inboard bearing house, (b) photo of the MFS with the PSS and accelerometers on the inboard bearing house",
        "texts": [
            " However, the strain caused by misalignment in the bearing house of the rotating machine can be reliably detected by the PSS due to its high sensitivity to strain [6]. It will also be demonstrated that the measurement of strain can help differentiate misalignment from other faults, such as unbalance, as the deformation of the bearing house is a primary effect of misalignment. To the best of the author\u2019s knowledge, this is the first work on the use of piezoelectric strain transducers for the detection of misalignment in rotating machinery. The experimental setup used in this work is shown in Fig. 1. The rotating machine used in the experiments is a SpectraQuest\u2122 machine fault simulator (MFS). The MFS is designed to study the vibration signature of common machine faults. The testing conditions are summarized in Table 1. Angular and parallel misalignment were introduced, and strain and vibration in the bearing house of the MFS were measured for different levels of misalignment, as well for rigid and flexible couplings used to connect the driving and driven shaft. Rotating the screws to displace the shaft rotor baseplate introduces the desired amount of angular, parallel or a combination of both misalignments in the MFS. As depicted in Fig. 1(a) and (b), strain is preferably measured in the bearing houses, as the primary forces caused by misalignment are transmitted to the other parts of the machine through these components. The PSS are placed on the top and bottom of the bearing houses, in order to measure the strains produced by radial and axial forces transmitted from the shaft to the bearing houses. The PSS on the top were placed close to the thinnest cross section of the bearing house, as thinner sections are more likely to deform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001537_0954406218791636-Figure17-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001537_0954406218791636-Figure17-1.png",
        "caption": "Figure 17. Representation of the turbine rotor-bearing-support system used in this example and its discretization for the implementation of TTMM. Additional masses have been incorporated to implement the rotating blades. Foundation is assumed rigid without permitting any motion. Bearing pad at each bearing is assumed to implement the total support mass (pedestal mass) and to perform motion, see Figure 7. Approximate slenderness ratio is L=D 7:3. Unbalance is applied within one case of single unbalance.",
        "texts": [
            " The risk of instability is high in such cases as the machine is supposed to retain its high rated speed, but with one (or both) bearing(s) to be lightly loaded. This is an excellent case to demonstrate how the adjustable bearing transforms its lemon bore configuration, into an offset halves configuration, ensuring the stability when partial arc steam admission occurs (yielding light bearing load), and offering the demanded design characteristics when normal operation is taking place (lemon bore configuration). In the following example, a small steam turbine, presented in Figure 17 and Table 7 with relative properties, performs a virtual run up in the time duration of 75 s from 0 to the rated speed of 7000 r/min and then it continuous for further 15 s at a constant rotating speed (rated speed) achieving a steady-state response where it is supposed to experience normal operation at full load, producing its rated power output. Then, at 90 s, the virtual partial arc steam admission occurs and a transient load is applied in the bearing #1 so as the resulting load to become 40% less than this of normal operating condition, while the turbine retains its rated speed of 7000 r/min"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001520_0954409718789531-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001520_0954409718789531-Figure6-1.png",
        "caption": "Figure 6. Bond graph sub-model of the wheelset (LW).",
        "texts": [
            " 1 2 rL rR\u00f0 \u00de \u00bc lyw \u00f052\u00de 1 2 rL \u00fe rR\u00f0 \u00de \u00bc r0 \u00f053\u00de 1 2 L R\u00f0 \u00de \u00bc 0 \u00f054\u00de 1 2 L \u00fe R\u00f0 \u00de \u00bc l \u00f055\u00de where, l is the conicity angle. Kinematic relationship. A wheelset is connected to a bogie frame through two primary suspensions. The velocity components of these attachment points in the corresponding directions are expressed by equations (56) and (57), which are used to obtain the transformer moduli required to develop the bond graph of the wheelset sub-system. These constraints are realized through 0-junction in the bond graph of the wheelset sub-system shown in Figure 6, which is denoted as LW. _Zwl _Ywl \" # \u00bc _Zw _Yw \" # \u00fe lp 0 0 0 _ c _ c \" # \u00f056\u00de _Zwr _Ywr \" # \u00bc _Zw _Yw \" # \u00fe lp 0 0 0 _ c _ c \" # \u00f047\u00de where _Ywl and _Ywr are the lateral and _Zwl and _Zwr are the vertical components of velocity of points, where the front left and right suspensions connect with the wheelset, respectively. The summation of right-hand side terms of equations (21) to (24) including the suspension force/ moment represents the resultant external force/ moment acting on the wheelset in/about the corresponding direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002312_1.4029054-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002312_1.4029054-Figure4-1.png",
        "caption": "Fig. 4 Clearance in situations with different support shapes",
        "texts": [
            " We determine this circle as initial clearance circle, namely, dashed circle depicted in Fig. 3. The radius of this initial clearance circle is Rc0 \u00bc Ro \u00fe q0 (19) where Ro is the radius of the outer race; and q0 is the initial protective clearance. The ACABD is designed to reduce the radius of the equivalent clearance circle. In order to analyze the influences of different supports on the reduction of this clearance circle\u2019s radius, supports with concave and convex shape surfaces are studied, as shown in Fig. 4. Suppose the section of support\u2019s surface is an arc with the center of curvature Om and radius Rm. In Fig. 4, point A denotes the nearest point on the arc to the outer race. The distance from point A to the outer race is q(hs); Os represents the center of revolute joint 1; l is the distance from Os to the arc\u2019s midpoint B; ls is the distance from Os to point A; and hs denotes the swing angle of the support. Suppose Rm of concave arc is negative and Rm of convex arc is positive. From Figs. 4(a) and 4(b), the real-time clearance q(hs) during the support rotation can be obtained: q hs\u00f0 \u00de \u00bc OdOm Rm Ro for concave shape support OdOm Rm Ro forconvex shape support ( (20) where OdOm \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0Rm l\u00de2 \u00fe OdOs 2 \u00fe 2\u00f0Rm l\u00deOdOs cos\u00f0hs\u00de q , OdOs \u00bc l\u00fe Ro \u00fe q0, Od represents the center of the ACABD",
            " Based on the force analysis, the approximate total torque acting on all supports can be gained: Ms \u00bc Ftlt Fnln Fsrs (25) Therefore, the dynamic equation for supports\u2019 swing can be obtained: NsIs \u00fe lq Rl Il \u20achs \u00bc Ftlt Fnln Fsrs (26) where Ns is the number of the supports; Is and Il mean the MOI of each support and the MOI of the link, respectively; Rl is the radius of the link, as shown in Fig. 5; rs represents the radius of revolute joint 1; lq denotes the distance between the centers of revolute joint 1 and 2; ln and lt represent the force arms of Fn and Ft at the contact point A relative to revolute joint 1, as shown in Fig. 4. For concave support, ln and lt are ln \u00bc \u00f0Rm l\u00deOdOs OdOm sin\u00f0hs\u00de lt \u00bc Rm \u00f0Rm l\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 OdOs sin\u00f0hs\u00de OdOm 2 s 8>>< >>: (27) As to convex support, ln and lt can be gained: ln \u00bc \u00f0Rm l\u00deOdOs OdOm sin\u00f0hs\u00de lt \u00bc Rm \u00f0Rm l\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 OdOs sin\u00f0hs\u00de OdOm 2 s 8>>< >>: (28) According to Coulomb\u2019s law of friction [27], Fs can be gained: Fs \u00bc ld ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F2 n \u00fe F2 t p _hs > 0 ls ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F2 n \u00fe F2 t p _hs \u00bc 0 ld ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi F2 n \u00fe F2 t p _hs < 0 8>>< >: (29) Generally, Ft\u00bc 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001155_s00202-017-0626-z-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001155_s00202-017-0626-z-Figure4-1.png",
        "caption": "Fig. 4 Topology of the test bench",
        "texts": [
            " load torque applied by dynamometer is equal to the external load torque reference Text. At low frequencies, the load machine\u2019s load torque resembles the value of Text, whereas at higher frequencies, tracking can no longer be assured. Therefore, this torque input is used mostly for the static load torque tests (e.g. step reference of load torque). The control structure for DEML has been experimentally verified on the test stand built with industrial components. The stand is shown in Fig. 3 and its topology in Fig. 4. The following devices were used: \u2022 SIMOTION D425 motion control system (1), \u2022 Smart Line Module 10 kW (2), \u2022 Double Motor Module 2 x 9A (3), \u2022 SITOP 24V DC Power Supply (4), \u2022 Sensor Module Cabinet SMC 20 (5), \u2022 SMPM machine, 3000 rpm, 7.3 Nm (6), \u2022 IM machine, 1455 rpm, 3 kW, 6,2 A (7), \u2022 PC with SCOUT and DCC software tools (8), \u2022 Flexible mechanical clutch (9). Induction machine is used as a DUT, controlled with the speed control loop. Permanent magnet synchronous machine is used as a dynamometer, controlled with the torque control loop"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002296_s12239-014-0093-8-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002296_s12239-014-0093-8-Figure1-1.png",
        "caption": "Figure 1. Right suspension mechanism 3D model.",
        "texts": [
            " Two important characteristics are proposed in connection with these structures: \u2022 the camber angle is constant in the case of the rigid mechanism and variable at the existence of deformable components; \u2022 the compensation of the rear wheel track variation determined trough actuator at rigid mechanism, this leads to an active track suspension mechanism with an operating advantage: elimination of the lateral forces on the wheel caused by this variation. The influence of these features highlights important conclusions for practice. A 3D view of the proposed mechanism is given in Figure 1 for the right wheel, it being similar for the other wheel. The kinematic scheme of this mechanism is drawn in Figure 2. (1) The two parallelogram mechanisms are the joints: J, C, D and I, and, respectively, E, F, G, H. (2) The actuator 1 moves the mobile frame 2 through the joint B. (3) In a plan representation the mobile frame is connected to the outer and inner arms 7 and 8 by the joint D and C. Also this mobile frame is connected to the lower and upper arms 3 and 6 by the joint E and H. The element 5 is a mobile plateau that makes the bump and rebound on vertical of the wheel",
            " (3) Patran as preprocessor gives a command to MD R2 Nastran that generates the MNF files for flexible elements (mobile frame 2 and upper 3 and lower 6 arms). (4) The STEP files for rigid components (others the ones deformable) and the MNF files for flexible elements are imported in ADAMS MD R2/View and jointed with rigid or compliant bushings. Then the motions of mechanism components are simulated considering the variable distance AB of the actuator depending on the bump or rebound (applied by a vertical oscillator, Figure 1) to the wheel. (5) Finally the variations of the two interesting quantities (camber angle and wheel track) are established by ADAMS post processing depending on the bump and rebound of the wheel for several cases of study. The study cases are: (1) rigid joints and arms (of reference); (2) compliant bushings and rigid arms; (3) rigid joints and flexible arms; (4) compliant bushings and flexible arms. The following data are chosen for the simulation process: \u2022 the force F in the wheel contact patch (Figure 1) is 0 N (case a), 2,000 N (case b), and 4,000 N (case c); \u2022 the translational bushing stiffness is 5,000 N/mm (conventional practice value and constant during the simulation process); \u2022 the material characteristics of steel for arms: Young\u2019s modulus E = 2.1\u2022105 MPa; Poisson coefficient \u03bd = 0.3; density \u03c1 = 7,800 kg/m\u00b3. The actuator displacement determined using the equations presented above was used to keep the wheel track constant in case 1 is also transmited to the mobile frame in the other cases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000405_s00170-015-6830-y-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000405_s00170-015-6830-y-Figure7-1.png",
        "caption": "Fig. 7 Analytical model of axial stresses of the bi-layered tube under internal pressure",
        "texts": [
            "8py), the inward defect disappeared, but there was a little bit of section ovalization. When the pressure reached the yielding pressure of 25 MPa (1.0py), the deformation of two-layer tubes showed a coherent mechanical behavior. No wrinkling, inward, and ovalization defects took place, as shown in Fig. 5f. Finally, sound tubes can be formed under the internal pressure of 25 MPa, as shown in Fig. 6. The reason for prevention of the wrinkling will be discussed next. An analytical model is established on the axial stress of bi-layered tube under the internal pressure, as shown in Fig. 7, where A-A section is the symmetry plane, \u03c3zi is the additional axial stress of the inner tube, and \u03c3zo is the additional axial stress of the outer tube. By the axial force equilibrium condition, we have \u03c3zi\u03c0dti \u00fe \u03c3zo\u03c0dto \u00bc \u03c0d2 4 p \u00f03\u00de Based on the deformation compatibility, the axial strains of the inner tube \u03b5zi and outer tube \u03b5zo should be equal at section A-A. Then by Hooke\u2019s law, \u03c3zi Ei \u00bc \u03c3zo Eo \u00f04\u00de Fig. 5 Wrinkling behavior and axial stress under different levels of internal pressure. a 0 MPa"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003839_acs.2982-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003839_acs.2982-Figure1-1.png",
        "caption": "FIGURE 1 Free-body diagram of the 3-DOF helicopter [Colour figure can be viewed at wileyonlinelibrary.com]",
        "texts": [
            " The forgetting factor l in (26) aims to guarantee the boundedness of the auxiliary matrices M and N. The constant \ud835\udf05 in (44) affects the convergence of the parameter estimation error W\u0303 . Finally, the selection of the adaptive learning gain \u0393 in (14) and (44) should be considered to trade-off the convergence rate and robustness, which is not necessary to be large to retain the parameter estimation convergence in this paper due to the use of leakage term H. In this section, the model of a benchmark 3-DOF helicopter (see Figure 1) is used as the numerical example to validate the proposed MRAC scheme. We refer to Quanser Inc.43 for more details on the modeling of this plant. In this study, we focus on the control of pitch motion of the helicopter for demonstration. The differential equation describing the pitch motion can be given as J\ud835\udf14?\u0308? = 2LaK\ud835\udc53Ku + (mwLw \u2212 (m\ud835\udc53 + mb)La)g cos (\ud835\udf14 + \ud835\udf140) , (51) where \ud835\udf14 denotes the pitch angle, J\ud835\udf14 is the moment of inertia, which is equal to (m\ud835\udc53 + mb)L2 a + mwL2 w, and the physical meaning of other parameters and their values can be found in Quanser Inc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003911_s42417-019-00111-6-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003911_s42417-019-00111-6-Figure4-1.png",
        "caption": "Fig. 4 Illustration of the bearing test rig",
        "texts": [
            "46\u00a0Hz and 158.46\u00a0Hz, respectively, and this result is very close to the result of Ref. [24]. The value in Ref. [24]., the extracted feature value, the calculated value and the error are given in Table\u00a02, it confirms that this proposed method can extract the fault feature accurately and effectively. Case 2: Detect the\u00a0Bearing Fault in\u00a0the\u00a0End\u2011to\u2011Failure The testing data of this case were obtained from Intelligent Maintenance System (IMS) in University of Cininnati [23]. The testing rig is shown in Fig.\u00a04, the four bearing Rexnord ZA-2115 were equipped on the shaft and driven by a motor. The motor rotating speed was 2000 r/min and radial load were 6000\u00a0lb. Every bearing was run to failure in this experiment. The dimension parameters of these tested bearings are given in Table\u00a03. There were a total of eight accelerometers 1 3 installed on the bearing housing. The sampling frequency of this experiment was 20\u00a0kHz and collected the vibration signal every 20\u00a0min. In this experiment, the outer race defect of bearing 1 was found until the 7th day"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001970_j.cja.2015.07.003-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001970_j.cja.2015.07.003-Figure1-1.png",
        "caption": "Fig. 1 An exploded view of a harmonic drive used in a spacecraft manipulator.",
        "texts": [
            " In general, compared with statistics-based models, physicsstatistics-based models assume in-depth understanding of the failure mechanism of corresponding components or materials, and are preferred whenever possible.17 However, since the failure mechanism and the performance degradation process of mechanical components are usually complex due to varying working conditions, an accurate physics-statistics-based ALT model cannot be easily established. A harmonic drive typically consists of three subcomponents: a wave generator (WG), a flex spline (FS), and a circular spline (CS), as shown in Fig. 1. The flex spline is a compliant element, and its deformation is essential for the operation of a harmonic drive.18 Generally, the flex spline of a harmonic drive exhibits a cup shape which is coaxially connected to the output shaft.19 To clearly show the circular spline, the flex spline is simplified as a gear-ring-shaped object in Fig. 1. A harmonic drive may work under different operating patterns via varying the transmission chains and the service environment. For instance, the circular spline can be driven by the wave generator and the fixed flex spline. Besides, torque can be transmitted out of the flex spline aided by the wave generator while locking the circular spline. No matter what operating pattern a harmonic drive is working in, the flex spline deforms periodically, and the influences of material and geometrical model on the performance of the harmonic drive in this periodical deforming process have been deeply studied"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000609_iemdc.2013.6556188-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000609_iemdc.2013.6556188-Figure7-1.png",
        "caption": "Fig. 7: Stator cross section",
        "texts": [
            " In testcase 2 for contrast the results of the k-shift approach are much more accurate than of the T-equivalent network (fig. 5(a)). Especially the k-shift model predicts the heat flux direction correctly (fig. 5(b)). The results are also improved against the classical approach with 2 nodes. This means that the compensation effect and the improvement in transient does not result from the higher discretization level of k-shift but from the node shifting technique. In this section it is shown how the k-shift approach can be applied to a two dimensional cross section of electrical machines stator (fig. 7). Symmetry is assumed in the middle of the tooth and the slot, the remaining geometry is separated in 8 control volumes 1-8. The classical lumped parameter approach and the Tequivalent network approach are shown in fig. 8(a) and fig. 8(c) respectively. The application of the k-shift approach is shown in fig. 8(b) and bases on a shift of the nodes starting from the classical lumped parameter approach. This is done without equipping the control volumes with a second node as derived above. This means that with this application of the k-shift model not every control volume is represented but a corresponding couple of control volumes respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001846_978-90-481-9707-1_116-Figure11.3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001846_978-90-481-9707-1_116-Figure11.3-1.png",
        "caption": "Fig. 11.3 Quadrotor arm and frame design in SolidWorks",
        "texts": [
            "5 details the total parts in the design of the quadrotor MAV, with specific weight budget assigned to each of them. On the right of the table listed the real weight of the components or designed parts, which will be detailed here and also in the next sections. The structure design can be separated into two different parts: 1. Quadrotor arms: A protection scheme is developed as a shell to encapsulate the motor into a tight chamber to fix its position. Carbon fiber beam can be mounted and screwed to the side, as shown in Fig. 11.3a. 2. Main body: The main frame structure has four slots for the quadrotor arms and contains two layers, where the avionic system is placed on the top and the battery is located in the lower level, as shown in Fig. 11.3b. It is fabricated with acrylonitrile butadiene styrene (ABS). The mechanical structure layout of the proposed quadrotor MAV is shown in Fig. 11.4a with all the parts assembled together as shown in Fig. 11.4b. Distance of two diagonal rotors is 142.54 mm, with a total height of 22.50 mm. Details of weight breakdown are reflected in Table 11.5. In this table, the estimated weight is approximated based on the design guideline shown in this chapter, while the current weight on the right of the table is the measured weight of the quadrotor MAV prototype, code name KayLion, made by National University of Singapore, following the guideline"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003805_s2301385019400053-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003805_s2301385019400053-Figure1-1.png",
        "caption": "Fig. 1. Coordinate system.",
        "texts": [
            " Ignore the curvature of the earth and the gravitational acceleration is constant; . Ignore the influence of the earth\u2019s rotation and revolution. Consider the ground coordinate system to be the inertial system; . Consider the quadrotor as a rigid body and the center of the quadrotors coincide with the origin of the body coordinate system; . Quadrotors are geometrically symmetrical and have a uniform mass distribution. The inertia matrix can be simplified to I \u00bc diag\u00bdIxx Iyy Izz . Ixx , Iyy , Izz represent the inertia corresponding to the three axes. As shown in Fig. 1, OeXeYeZe denotes an earth-fixed inertial frame and ObXbYbZb a body-fixed frame. The orientation of the rigid body is given by a rotation matrix Rb e , and we will U n. S ys . 2 01 9. 07 :4 7- 54 . D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by T H E U N IV E R SI T Y O F N E W S O U T H W A L E S L IB R A R Y o n 08 /2 5/ 19 . R eus e an d di st ri bu tio n is s tr ic tly n ot p er m itt ed , e xc ep t f or O pe n A cc es s ar tic le s. use Z Y X Euler angles to model this rotation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000524_j.jsv.2014.05.041-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000524_j.jsv.2014.05.041-Figure1-1.png",
        "caption": "Fig. 1. Rotational degree-of-freedom model of a single-mesh gear system.",
        "texts": [
            " Two different speed fluctuation models are considered. Finally, the influences of parameters on instabilities are systematically investigated, such as the amplitude of mesh stiffness variation, the characteristics of speed fluctuations and damping. The single-mesh gear dynamic model used is based on the typical linear time-varying model developed by Sika and Velex [17]. Translational degrees of freedom are ignored, and only rotational vibrations \u03b81 and \u03b82 relative to the rigid body gear rotations are considered (Fig. 1). In general, the input speed for rigid-body conditions can be introduced via Fourier series as \u03a9\u00f0t\u00de \u00bc\u03a90\u00fe\u03a9a sin \u00f0\u03c9at\u00fe\u03c8\u00de (1) Generally, the gear mesh stiffness is expressed as Fourier series. In order to obtain analytical solutions, only the firstorder harmonic component is taken to simplify the analysis. This processing will cause an error because higher-order components are ignored. However, it does not affect the analysis process and the phenomenon presented in the gear system. More accurate forms of mesh stiffness are available in [19\u201321]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002170_02640414.2014.1003586-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002170_02640414.2014.1003586-Figure1-1.png",
        "caption": "Figure 1. 3D link segment model of a cricket bowler depicting the initial point within the initial segment (I1), segments within the chain (I1 \u2013 I10), global angle of the initial segment (\u03b81), intermediate joint angles between the initial segment and the endpoint.",
        "texts": [
            " Given the use of linear statistics in the previous literature can only be viewed as an approximation of the relationship between elbow kinematics and ball release velocity, a simulation model that accounts for the inherently non-linearity of the relationship between rotational execution variables and outcome variables (Craig, 1989) may be more appropriate. The position of the endpoint of a multi-joint kinematic chain is a direct function of the location and orientation of the initial point within the chain, the dimensions of the segments within the chain and the intermediate joint angles between the initial segment and the endpoint of the chain (Zatsiorsky, 1998, Figure 1). The linear velocity of the endpoint can then be obtained through numerical differentiation of these variables. Such a model, termed in this study a \u201cForward Kinematic Model\u201d (FKM), is well suited to assessing the relationships between individual or composite execution variables and endpoint velocity, as the non-linearity of the system is modelled appropriately. Furthermore, the use of a FKM will allow the effects of alterations of a bowler\u2019s elbow joint kinematics on wrist joint velocity to be determined"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003778_1.4042636-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003778_1.4042636-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of 4-UPS-UPU parallel mechanism",
        "texts": [
            " 2, the kinematic model of spherical clearance is established according to the kinematic relation between the spherical joint elements. In Secs. 3, the contact force model of spherical clearance, which contains normal contact force model and tangential friction force model, is established. In Sec. 4, the dynamic equation of the mechanism that considered spherical joint clearance is derived based on Newton\u2013Euler method. In Sec. 5, the dynamic response with different clearance values and nonlinear characteristics of the parallel mechanism are analyzed. 2.1 Eccentricity Between Joint Elements With Clearance. As shown in Fig. 2, the fixed coordinate system OA XAYAZA denoted as fAg is set up on the fixed platform of 4-UPS-UPU parallel mechanism. The moving coordinate system OB XBYBZB denoted as fBg is set up on the moving platform. The branch chain coordinate system Ci XCiYCiZCi denoted by fCig is set up on the driving limbs, where i\u00bc 1, 2, 3, 4, and 5. When the clearances are not considered, the center of the sphere is coincident with the center of the ball sleeve and the degree-of-freedom is 3. It is assumed that there is a clearance at the spherical joint that between the third driving limb and the moving platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002173_detc2014-35099-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002173_detc2014-35099-Figure2-1.png",
        "caption": "Figure 2 Surface issue (a) chordal error (b) staircase effect [1]",
        "texts": [
            " In order to present a generic view of inherent process errors, this section describes the error generation issues and their reduction strategies and reviews different key issues involved in the part quality improvements. There are two common approaches used to control surface quality: chordal error reduction is the first approach and staircase error reduction is the second approach. No matter which defect is reduced, non-smooth surface is relatively much more visible if curvilinear surfaces are used instead of planar surfaces. During the tessellation of the geometric model of desired parts, triangles are used to approximate the boundary of the CAD model. Fig. 2 (a) shows the definition of chordal error, caused due to tessellation of geometric model. Chordal error is a surface defect that can be reduced to some extent by increasing number of triangles during tessellation of CAD model of the part. However, increasing number of triangles lead to computational complexities. Furthermore, the original curvature of a surface is non-recoverable by this operation. Hence, it does only sub optimization of surface quality. Second surface defect is a generic limitation of the layered manufacturing building process and is known as staircase error and staircase error minimization is one of the effective 1 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use techniques of surface quality improvement [1]. It is induced in LM build parts due to layer by layer nature of the process as shown in Fig. 2(b). As the layer thickness is used to approximate the curved surfaces in LM processes the selection of a proper layer thickness of the part during the building process can reduce the stair-steps and bring surface roughness within acceptable limits. Hence, selection of a proper layer thickness can better approximate the given curves. Therefore, the layer thickness at which the part is built has a significant effect on the quality of the various surfaces of the part. The layer thickness of the part also affects other factors such as volumetric error, build time, accuracy etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001934_1.4030344-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001934_1.4030344-Figure9-1.png",
        "caption": "Fig. 9 The (a) radial and (b) perpendicular polarized magnetic bearings",
        "texts": [
            "org/about-asme/terms-of-use where Fy;s;i \u00bc r1r2 4pl0 R za; i\u00f0 \u00de \u00fe R za \u00fe H B; i\u00f0 \u00de \u00fe R za \u00fe H; i\u00f0 \u00de \u00fe R za B; i\u00f0 \u00de\u00f0 \u00de and R a; i\u00f0 \u00de \u00bc \u00f0W 2 \u00fe i 1\u00f0 \u00de2p n W 2 \u00fe i 1\u00f0 \u00de2p n \u00f02p 0 \u00f0R4 R3 \u00f0R2 R1 e\u00fe r12 cos h\u00f0 \u00de r34 cos h0\u00f0 \u00de\u00f0 \u00der12r34 r2 12 \u00fe r2 34 \u00fe e2 2r12r34 cos\u00f0h h0\u00de \u00fe 2e r12 cos h\u00f0 \u00de r34 cos h0\u00f0 \u00de\u00f0 \u00de \u00fe \u00f0a\u00de2 1:5 dr12dr34dhdh0 Mathematical Modeling of RMD Magnetic Bearing. To enhance the load carrying capacity of magnetic bearing, an RMD configuration as shown in Fig. 8 has been proposed. The configurations have two axial magnets and one radial magnet in rotor as well as in stator. In Fig. 8, H is the length of stator and B is the length of rotor. From this figure, it can be observed that for developing theoretical formulation for the RMD configuration, the load carried by the repulsion force between two radially polarized magnets (as shown in Fig. 9(a)) and perpendicular polarized magnets (as shown in Fig. 9(b)) is to be formulated. The total load carrying capacity of RMD configuration (Fy,RMD) is given by Fy;RMD \u00bc XhX \u00f0Radial polarized\u00de \u00fe X \u00f0Axial polarized\u00de \u00fe X \u00f0Perpendicular polarized\u00de i (8) Yonnet [17] proved that for an identical bearing, the load carrying capacity between two radial polarized magnets remains same as load capacity between two axial polarized magnets. Hence, a separate formulation for radial polarized magnets is not required and Eq. (8) can be modified for the present case as Fy;RMD \u00bc Xh 3 X Axial polarized\u00f0 \u00de \u00fe 4 X Perpendicular polarized\u00f0 \u00de i (9) A separate modeling for perpendicularly polarized magnets is required, formulation of which is described in the following subsection. Mathematical Modeling for Perpendicularly Polarized Magnets. In perpendicularly polarized magnetic bearing, the polarization direction between rotor and stator is 90 deg to each other. For example, in Fig. 9(b) the rotor is radially polarized and stator is axially polarized. The vertical force between the two rings in a perpendicularly polarized magnetic bearing (Fy,p) is represented as Fy;p \u00bc r1r2 4pl0 A za;R1\u00f0 \u00de A za\u00feB;R1\u00f0 \u00de A za;R2\u00f0 \u00de\u00feA za\u00feB;R2\u00f0 \u00de\u00f0 \u00de (10) where A za;R1 \u00f0 \u00de \u00bc \u00f02p 0 \u00f02p 0 \u00f0H 0 \u00f0R4 R3 e\u00fe R1 cos h\u00f0 \u00de r34 cos h0\u00f0 \u00de\u00f0 \u00de R2 1 \u00fe r2 34 \u00fe e2 2R1r34 cos\u00f0h h0\u00de \u00fe 2e R1 cos h\u00f0 \u00de r34 cos h0\u00f0 \u00de\u00f0 \u00de \u00fe \u00f0za z34\u00de2 dr34dzabdhdh0 For n number of sector magnets, Eq. (8) is modified and is given by Eq. (11) Journal of Tribology OCTOBER 2015, Vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003410_j.matpr.2017.11.530-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003410_j.matpr.2017.11.530-Figure5-1.png",
        "caption": "Fig. 5: simple machining model as input in DEFORM 3D software",
        "texts": [
            " In this study, DEFORM 3D was used (this software has been exclusively built with machining operations in mind, thus being optimized for metal cutting operations). It has a user friendly interface and can simulate both 2D and 3D machining operations with a wide range of materials in the property library. The software uses adaptive meshing to increase the accuracy of the simulation. Although the user can\u2019t configure the controls of the solver, DEFORM is a good software that allows fast machining setups because of its simple and clean interface as shown in figure 5 [9]. The finite element mesh was generated using DEFORM automatic mesh generation system. Re meshing parameters, including minimum element size, and parameters for absolute mesh definition are set within the system can be shown in figure 6 and 7. For these simulations, a minimum element size of 0.02mm was specified Sreeramulu et al., / Materials Today: Proceedings 5 (2018) 8364\u20138373 8369 Fig.6: Meshed Tool and Work piece Fig. 7: Mesh design of a work piece Heat exchange is defined on the boundaries A-D and D-C"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002106_j.apacoust.2014.04.001-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002106_j.apacoust.2014.04.001-Figure2-1.png",
        "caption": "Fig. 2. Schematic diagram of the experiment.",
        "texts": [
            " In order to obtain the multi-source noise acoustic signal with Doppler Shift, four speakers are mounted at different locations in the car as shown in Fig. 1. Four speakers would play different acoustic signals while the car passes by the microphone, which is placed about 1.7 m away from the running track of the car, along a straight line with a constant speed. During the pass-by test, the microphone receives the multi-source acoustic signal with Doppler Shift and then recorded by data acquisition device. The schematic diagram of the experiment is shown in Fig. 2, where the notations and parameters are explained as follows: S is the location of signal source (represents the train bearing); A, B and C represent locations of noise sources and the spacing between the signal source and each noise source is: dsa = 2.46 m, dsb = 1.56 m, dsc = 2.30 m; the microphone and the photoelectric sensor are placed parallel with the vector of the car\u2019s velocity and the distance between them is rs = 1.7 m; the height of the microphone is equal to that of S; xs = 4 m is the distance between the microphone and the photoelectric sensor; the distance between two wheels is dw = 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001373_j.triboint.2018.01.034-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001373_j.triboint.2018.01.034-Figure5-1.png",
        "caption": "Fig. 5. Driven disk tested in the experiment. (a) 2D model. (b) Installation diagram of pressure sensors in the driven disk.",
        "texts": [
            " torque, is measured by the torque-speed sensor while the film thickness is decreasing over a reference command pressure. Since both of the disks are rotating, A displacement sensor (linear measurement range: 0\u20135mm) is mounted on the closed tank to obtain the film thickness by indirect measurement. The test conditions for measuring the flow of the oil film are shown in Table 2. During the tests, the effect of axial deformation and radial deformation on dynamic characteristics is carried out to verify the developed model. No. 8 ATF acts as a lubricant and a coolant. Fig. 5(a) shows the 2D model of the driven disk tested. The related parameters are listed in Table 3. It is speculated that #1 specimen can be used in the test in which the disks make up parallel interface. And there are #2~#4 specimens to be used as the interface of axial deformation. Similarly, the uses of #5~#6 specimens are related to the test about the effect of radial deformation. In order to facilitate installation of pressure sensors during the measurement of oil pressure, the driven disk was fixed in the test bench by maximizing the excitation current of magnetic powder brake. The probes were installed in the driven disk as shown in Fig. 5(b) in which the areas of section line represent the grooved areas. P1, P2, P3 denote pressure sampling points in the flow respectively. The oil pressure was measured by means of dynamic pressure transducers(probe diameter: 5 mm, measurement range: 0\u20132MPa). T1, T2, T3 represent temperature sampling points in the flow respectively. Temperature was measured with platinum resistance temperature sensor Pt100 (probe diameter: 3 mm,measurement range: 0\u2013100 C). The Pt100was mounted in Teflon insert which was screwed into the driven disk so as to reduce the impact of temperature on experimental results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002124_s12010-014-0988-x-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002124_s12010-014-0988-x-Figure1-1.png",
        "caption": "Fig. 1 Schematic illustration of a cross-sectional view of polymerized electrode",
        "texts": [
            " A 50-mM phosphate buffer solution (PBS) of pH 6.8 was prepared using NaH2PO4, Na2HPO4, and deionized (DI) water. 2. Electrochemical cell All electrochemical experiments including activity assay, electropolymerization, and response current measurements were performed at room temperature in a three-electrode electrochemical cell using electrochemical station (Autolab PGSTAT302N). Platinum foil and Ag/AgCl were used as counter and reference electrodes, respectively. 3. Fabrication and characterization of electrodes Figure 1 presents the schematic illustration of cross-sectional view of fabricated PPy/Pt/ Anodisc\u2122 electrode. Patterned Pt electrode (defined by mask) was coated on the Anodisc\u2122 using direct current magnetron-sputtering system operated at base pressure of 2\u00d710\u22127 mbar. During deposition, argon gas flow rate was maintained at 20 cm3 min\u22121. Deposition rate and pressure during sputtering were 1 \u00c5 s\u22121 and 5\u00d710\u22123 mbar, respectively. The substrates were rotated at a speed of about 30 rotations per minute (rpm) during sputtering in order to ensure uniformity of a deposited metal film",
            " Pt/Anodisc\u2122 was used as a working electrode for electrodeposition. Aqueous solution was prepared containing freshly distilled pyrrole monomer (25 mM) and LiClO4 (100 mM) as a supporting electrolyte. Prior to electropolymerization, solution was deaerated by bubbling nitrogen gas for 10 min. Electropolymerization was carried out potentiostatically at 1.8 V applied potential. Optimization of thicknesses of Pt and PPy nanotubes is of utmost importance in order to obtain the nanoporous electrode as shown in Fig. 1, for the improved performance of biosensor. Hence, Pt thickness was varied from 25 to 100 nm, and polymerization duration was varied from 10 to 100 s for different samples. The surface morphologies of the produced Pt and PPy films were investigated using field emission scanning electron microscope (FESEM, Carl Zeiss SUPRA55). Also, porosity of Pt-coated electrode before and after polymerization for different durations was calculated using an image processing tool in Matlab by transforming the grayscale images obtained from FESEM into binary images (having only black and white regions, where black regions correspond to the regions containing pores on the electrode), and thus, a calculation based on the ratio of area occupied by black pixels to the total area of the image (black+white) led to the estimation of the porosity of the Pt/PPy-modified Anodisc-based electrodes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003170_0954408917727198-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003170_0954408917727198-Figure3-1.png",
        "caption": "Figure 3. Schematic diagram of experimental set-up.",
        "texts": [
            " Based on the presented TCC model, the dimensionless parameters were introduced for the purpose of the results generalization23 F \u00bc F AaE , H \u00bc Ht lA 1 2 a , A r \u00bc Ar Aa \u00f027\u00de In order to verify the presented TCC model, an experimental set-up with annular interface was designed in this paper. The one-dimensional steady measurement method was introduced to obtain the TCC of the annular interface, where the radial steady gradient of the heat flow in one-dimension for the testing pieces could be measured directly.16 The schematic diagram of the experimental set-up is presented in Figure 3. To ensure the heat flow spread radially, a heating device was arranged in the middle of the entire device and a cooling device was arranged externally to the outer testing piece. The heat insulating materials were placed in the upper and lower parts of the test pieces and the shielding was utilized to form a vacuum confined space for the air thermal convection effect reduction. The PT100-type temperature sensors were arranged on the same plane to reduce the measurement error. Ri,j are the distribution radii of the temperature sensors, where the subscript i \u00bc 1, 2, 3 indicated the calibrated copper ring, the inner testing piece and the outer testing piece, j \u00bc 1, 2 indicated the number of temperature sensors in each test piece"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure9.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure9.1-1.png",
        "caption": "Fig. 9.1 To determination of the impedance of a single-cage rotor slot bar (a) simplified scheme of the single-cage rotor slot; (b) the single-cage rotor slot as a layered structure",
        "texts": [
            " Because of the symmetry of the slot arrangement, we can use the rotor area enclosed within a single tooth division. We assume that the permeability of the rotor iron is a relatively large value, i.\u0435., \u03bcF 1. The fields distributed in the teeth and rotor joke regions are then absent, and thus the rotor field can be considered localized in a single slot. We will assume that the rotor slot is magnetized by a purely tangential field. In accordance with the system of orientation of coordinate axes accepted in Fig. 9.1, the magnetic field strength in the rotor slot will have only the x \u2013 component. The electric field strength is represented only by the z \u2013 component. The vectors of the electric and magnetic field strength are changed in the rotor slot only in the direction of the y axis. In reality, the picture of the rotor slot field is different from that accepted in Fig. 9.1. Moreover, this difference does not significantly distort the results of the calculations. Therefore, the consideration of the real picture of the slot field can be implemented using the special factors in the final expressions for the rotor slot bar impedance. We now consider the rectangular slot in which the conducting rectangular winding bar with height h\u03a0 is located. As shown in Fig. 9.1, in such slot, the magnetic lines are directed perpendicular to the surfaces of the teeth walls, and they are distributed parallel to the bottom of the slot. This means that a one-dimensional field characterized by the conditions Hz\u00bc 0 and Hy\u00bc 0, and Ex\u00bc 0 and Ey\u00bc 0 takes place in the rotor slot. Taking into account these conditions, and also bearing in mind that Ez\u00bcE, the field equations given in (7.33) can be represented as rotH \u00bc dHx dy \u00bc E \u03c12 ; rotE \u00bc dE dy \u00bc j\u03c91s\u03bc0Hx; \u03c3 \u00bc E \u03c12 \u00f09:7\u00de 332 9 The Single-Cage Rotor: The Slot Leakage Circuit Loops In (9.7), the first equation we differentiate with respect to y, and then using the first and second equations, for the magnetic field strength in the rotor slot we can have dH2 x dy2 \u00bc j\u03c91s\u03bc0 1 \u03c12 Hx \u00bc \u03b222Hx \u00f09:8\u00de where \u03b22 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi j\u03c91s\u03bc0 1=\u03c12\u00f0 \u00de p An analogous equation can be obtained for vector E. The picture of the field distribution in the rotor slot is known (Fig. 9.1). Therefore, equation (9.8) can be also obtained using the laws of total current and electromagnetic induction. Here, a sufficiently high level of clarity with respect to the technique for obtaining equation (9.8) is provided. To use the law of total current, we separate the contour with height dy and width b\u03a0 on the surface of Fig. 9.1. The equation for this contour using the law of total current then takes the form Hxb\u03a0 Hx \u00fe \u2202Hx \u2202y dy b\u03a0 \u00bc E \u03c12 b\u03a0dy From here, we have \u2202Hx \u2202y \u00bc E \u03c12 \u00f09:9\u00de On the surface perpendicular to the field lines (Fig. 9.1), we separate the contour with height dy and length l. Applying the law of electromagnetic induction to this contour, we can obtain El\u00fe E\u00fe \u2202E \u2202y dy l \u00bc \u03bc0 \u2202Hx \u2202t ldy or \u2202E \u2202y \u00bc \u03bc0 \u2202Hx \u2202t \u00f09:10\u00de The equation shown in (9.9) we differentiate with respect to y. Then, using equations (9.9) and (9.10), we can show that the equation for the magnetic field strength takes the form of (9.8). The solution of equation (9.8) can be represented as Hx \u00bc Ae\u03b22y \u00fe Be \u03b22y \u00f09:11\u00de 9.1 The Single-Cage Rotor: Slot Leakage Single Circuit Loop 333 From (9",
            "1 The Single-Cage Rotor: Slot Leakage Single Circuit Loop 335 where Zcz \u00bc j\u03c91\u03bc0 \u03b22 tcz b\u03a0 \u03beLcz \u03be\u03c4cz 1\u00fe e 2\u03b22h\u03a0 1 e 2\u03b22h\u03a0 The rotor slot leakage impedance Zcz can be reduced to the form Zcz \u00bc j\u03c91\u03bc0 \u03b22 tcz b\u03a0 \u03beLcz \u03be\u03c4cz 1\u00fe e 2\u03b22h\u03a0 1 e 2\u03b22h\u03a0 \u00bc j\u03c91\u03bc0 \u03b22th\u03b22h\u03a0 tcz b\u03a0 \u03beLcz \u03be\u03c4cz \u00bc \u03c12 h\u03a0s tcz b\u03a0 \u03beLcz \u03be\u03c4cz \u03b22h\u03a0 th\u03b22h\u03a0 \u00f09:19b\u00de On the basis of expression (9.19b), it can be shown that the resistance (rc2/s) and leakage reactance x\u03a02\u03c3 take the form of formula (9.6) obtained by the slot model using electromagnetic parameters \u03c1cz and \u03bcxcz determined from (9.1). From here, it follows that in relation to the average values of the field strengths, the slot bar can be represented as a conditional layer with thickness h\u03a0 and width tcz equal to the length of a single tooth division (Fig. 9.1b). In equations (9.16) and (9.17), the average values of field strength are represented by the vectors Ecz1,Hcz1 and Ecz2,Hcz2. In relation to vectors Ecz1, Hcz1 and Ecz2,Hcz2, these equations can be written in the form of four-terminal equations (for example, [1\u201314, 15, 16]). We can have Ecz1 \u00bc j\u03c91\u03bc0 \u03b22th\u03b22h\u03a0 tcz b\u03a0 \u03beLcz \u03be\u03c4cz Hcz1 \u00fe j\u03c91\u03bc0 \u03b22sh\u03b22h\u03a0 tcz b\u03a0 \u03beLcz \u03be\u03c4cz Hcz2 \u00bc z11Hcz1 \u00fe z21Hcz2 Ecz2 \u00bc j\u03c91\u03bc0 \u03b22sh\u03b22h\u03a0 tcz b\u03a0 \u03beLcz \u03be\u03c4cz Hcz1 \u00fe j\u03c91\u03bc0 \u03b22th\u03b22h\u03a0 tcz b\u03a0 \u03beLcz \u03be\u03c4cz Hcz2 \u00bc z12Hcz1 \u00fe z22Hcz2 \u00f09:20\u00de where z11 \u00bc z22 \u00bc j\u03c91\u03bc0 \u03b22th\u03b22h\u03a0 tcz b\u03a0 \u03beLcz \u03be\u03c4cz ; z12 \u00bc z21 \u00bc j\u03c91\u03bc0 \u03b22sh\u03b22h\u03a0 tcz b\u03a0 \u03beLcz \u03be\u03c4cz Equations (9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003620_j.matpr.2018.06.241-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003620_j.matpr.2018.06.241-Figure4-1.png",
        "caption": "Fig. 4 Bicycle subjected to condition",
        "texts": [
            " (ii)Steady State pedalling- the cyclist is seated on the bicycle and applying a force of 200N due to leg dynamics. The load is assumed to be concentrated at the bearing as shown in above Fig 3 (i -ii). (iii)Vertical Impact- vertical impact loads are represented by multiplying the cyclist\u2019s weight by some amount of G factor. In this case a factor of 2G is taken taking the load to 1400N which is the necessary case when an object falls from an infinitesimal height onto a rigid surface as shown in below Fig 4 and Fig 5. (iv)The case that is presented now is from [2]. Here the loads are simulated for the load bumps occurring at the front wheel. A resultant load of 2700N is transmitted at the rigid links which are then connected to the axle and then to the frame via the fork. 18924 Devaiah B.B et al./ Materials Today: Proceedings 5 (2018) 18920\u201318926 The frame is divided into 169770 triangular elements. The stress results for the different loading conditions are shown below for (i) (ii) and (iii) cases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure11-1.png",
        "caption": "Figure 11: Strut uparching and stress step-up in flux of force optimized lattice structure",
        "texts": [
            " This leads to the conclusion, that the bending loads in this sectional view have severely been reduced. Nevertheless, the maximum value for the von-Mises stress has risen to 74.63 N/mm\u00b2 (see figure 10 a). Under the assumption, that this maximum stress linearly depends on the applied force, a maximum load of 221.1 N can be calculated for the obtaining of the limit of elasticity. This means a worsening of 46 % compared to the periodic structure. The reason for the increased maximum stress can be seen in figure 11. The figure shows the sectional view at cutting plane 1 (see figure 3). It can be recognized, that there appears severe uparching, especially in longer struts with strong curvature. This uparching leads to high stresses along the struts and especially to stress peaks near the nodes. Although the ratio of the force at the limit of elasticity to the structure\u2019s mass (2.65 N/g) is just slightly below the value for the periodic structure (2.76 N/g) due to the reduced mass, this result is not satisfying",
            " This means an improvement of 7 % compared to the curved structure and 9 % compared to the periodic geometry. This leads to a value of 27.7 N/(mm*g) for the stiffness to mass ratio. Hence, an additional increase of 7 % results in contrast to the curved structure. Maximum Stress. Way more important than the enhancement of the stiffness is a decrease of the maximum of the appearing von-Mises stress compared to the curved structure. Figure 14 shows the sectional view at cutting plane 1. T In order to make the results more comparable, the colour scale has been chosen similar to figure 11. It is recognizable, that no or only marginal uparching appears in the single struts. This leads to significantly reduced stress states and stress peaks. The result is a maximum von-Mises stress of 31.9 N/mm\u00b2. Under the assumption, that this maximum stress linearly depends on the applied force, a maximum load of 516.8 N can be calculated for the obtaining of the limit of elasticity. This is an enhancement of 27 % compared to the periodic structure and 134 % compared to the flux of force adapted structure with curved struts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure10-1.png",
        "caption": "Fig. 10. Spherical body in the water flow.",
        "texts": [],
        "surrounding_texts": [
            "The drag coefficient is calculated by a classical formula [7]: Cd \u00bc 2 Fd q A U2 ; \u00f05\u00de where Fd \u2013 the drag force, obtained from simulation results, q \u2013 water density, U \u2013 mean velocity of a water flow, m/s, A \u2013 reference area, m2."
        ]
    },
    {
        "image_filename": "designv11_13_0003116_s11465-017-0452-z-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003116_s11465-017-0452-z-Figure10-1.png",
        "caption": "Fig. 10 Sensor locations on the WT (A, B, and C represent the main bearing, planetary gearbox, and generator, respectively) 1\u2013X direction of the gearbox input end; 2\u2013X direction of the first-stage planetary gear train; 3\u2013X direction of the second-stage planetary gear train; 4\u2013X direction of the fixed-shaft gear train; 5\u2013Y direction of the gearbox output end; 6\u2013Y direction of the generator input shaft; 7\u2013Z direction of the parallel gear train; 8\u2013X direction of the generator input end; 9\u2013X direction of the generator",
        "texts": [
            " The transmission ratios and orders of the transmission chain at each stage can be calculated using the structural parameters presented in Table 1. 4.2 Vibration signal acquisition The vibration signals collected from the surface of the WT contain rich information about the health condition of the WT [63]; therefore, vibration analysis was performed in this study for WT fault signature extraction. The vibration signals of the investigated WT were collected by accelerometers mounted on the planetary gearbox housing and generator housing in X, Y, and Z directions, as shown in Fig. 10. The unit of the collected vibration signal in each channel is m/s2. 4.3 Validation of the proposed technique for WT tacholess order tracking The vibration signals sampled from WTs based on the presented experimental setup were subjected to further signal processing to validate the effectiveness of the proposed method. The vibration signals were continuously collected with accelerometers at a sampling frequency of 25.6 kHz and a sampling length of 4 s. Two sets of WTs numbered 035# and 049# were measured"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002626_12.2222130-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002626_12.2222130-Figure3-1.png",
        "caption": "Figure 3. 3D-printed ABS plastic cube without damage, orthographic view.",
        "texts": [
            " First, the program instructs the DAQ to generate a sinusoidal excitation at 80.5 KHz which is then amplified by the amplifier with a gain setting of 20. Second, the laser Doppler vibrometer is directed to scan across a defined scan region and directed by the two Galvo steering mirrors. The magnitude and phase of the 80.5 KHz response is estimated at each pixel to provide a full-field, steady state response (Figure 5). In this research, a 6 cm long by 2 cm wide by 1 cm tall cuboid was printed using ABS plastic, as pictured in Figure 3. An initial \u201chealthy\u201d print without any imposed damage was created as an experimental control. The second print measured the effects of foreign object damage (FOD) on the composition and structural properties of the 3D-printed cube. FOD was created through the addition of triangle-shaped section of electrical tape, approximately 1 cm by 1 cm, to the top of the unfinished part, specifically during the printing of the 27th layer, as shown in Figure 4 (left). Proc. of SPIE Vol. 9804 980418-3 Downloaded From: http://proceedings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001934_1.4030344-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001934_1.4030344-Figure12-1.png",
        "caption": "Fig. 12 Rotor for RMD configuration: (a) rotor and (b) rotor structure with square magnet",
        "texts": [
            " In the present work, an attempt has been made to demonstrate the enhancement of the static as well dynamic performance of the RMD configuration using the available magnets. In the case of RMD configuration, i.e., configuration 3, the axial and radial polarizations can be achieved by rotating the square magnets by 90 deg about the axis of rotation. Due to unavailability of desired size of the magnets in the market, the rotor for the RMD configuration has been developed using the available magnet. Magnets used in configurations 1 and 2 were full ring axially polarized magnets but due to unavailability of radial magnets for rotor, a structure similar to stator (Fig. 12(a)) was developed for the rotor as shown in Fig. 12(b) and the rotor developed for the RMD configuration is shown in Fig. 12(c). Figure 13 shows the RMD configuration and seven different arrangements possible in the RMD configuration. In \u201carrangement 1,\u201d the force is exerted between two axially polarized magnets with an axial offset of [(H2 H1) (B2 B1)]. The force exerted between the magnets can be estimated using Eq. (7). Similarly in the \u201carrangement 2,\u201d the force exerted between an axial polarized stator magnet and radial polarized rotor magnet can be estimated using Eq. (11). In \u201carrangement 3,\u201d the force is exerted between two radial polarized magnets which can be estimated using Eq",
            " Same way, the load carrying capacity of other arrangements Journal of Tribology OCTOBER 2015, Vol. 137 / 042201-7 Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use can be calculated and summed to find the total load carrying capacity. The advantage of the RMD configuration compared to backto-back configuration is that in RMD configuration, the axial and radial magnets are in partial attraction and hence does not require a special clamp or glue to hold them (which can be depicted from Fig. 12(b)) as it is required in the case of back-toback configuration. The measured flux density along the axial length of stators of full ring magnet and RMD configuration is shown in Fig. 14(a). The flux density was measured using Gauss meter as shown in Fig. 14(b). From these figures, it can be concluded that the developed RMD configuration is able to generate more magnetic flux compared to full ring magnets. The displacements of the shaft for different loads are shown in Fig. 14(c). From this figure, it can be inferred that the maximum load carrying capacity of full ring stator is higher than stator with square magnets, but the load carrying capacity is enhanced by using RMD configuration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001447_j.optlastec.2018.05.031-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001447_j.optlastec.2018.05.031-Figure13-1.png",
        "caption": "Fig. 13. Samples before tensile test.",
        "texts": [
            " The cross sections were prepared for microstructure analysis, the reagent CuSO4 20 g, H2O 80 ml, H2SO4 5 ml, HCl 100 ml was used. The samples were immerged in the etching solution and then polished with a suspension of 9 lm. Study of the microstructure was conducted using the microscope Axiovert Observer.D1m of \u2018\u2018Carl Zeiss\u201d, metric measurements were carried out using Thixomet image analyzer. The testing machine Shimadzu AG-5kNX was used for mechanical testing (UTS); microhardness was measured by manual equipment Remet HX 1000 at a load of 100 gr. After welding the samples were cut 5 mm wide using EDM Fig. 13(a). Tensile testing was carried out under controlled rate of load with deformation rate 0.5 mm per minute due to very small dimension of samples. When performing the experiment according to the plume shape, it can be assumed that at 900 W the welding was in the mode of heat convection laser welding (HCLW) and at 1200 W \u2013 keyhole laser welding (KHLW), it is confirmed by the microstructure. Despite the small difference in power ( 30%), the penetration depth measured by cross sections is significantly different ( 3\u20134 times, Table 4)",
            " In Figs. 11 and 12 and Table 6 the influence of heat treatment on the microhardness of the welding seam metal and the SLM base metal can be seen. The HT1 sample has the lowest microhardness and HT2 sample microhardness is slightly higher than that of HT1 sample, HT3 sample has the highest microhardness and NHT sample microhardness is slightly lower than that of NH3 sample. For carrying out mechanical tests, three samples were prepared for each heat treatment regime. Samples were cut by EDM cutting (Fig. 13a). In view of non-standard samples and accordingly nonstandard tests, not all tests were carried out till the sample was broken. It was also difficult to measure the cross-sectional area for all samples due to non-melting (Fig. 6), so the obtained strength values are approximate. Fig. 14 shows the samples after the tests; in all cases the deformation of CR was observed. In the case of HT1, a rupture in the welding seam occurred at all samples at a rela- tively low load ( 600 MPa). In the remaining cases, destruction did not occur, as the samples on the CR side deformed and got out of the grippers of the bursting machine, and this despite the non-melting in all cases (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001621_rpj-04-2017-0057-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001621_rpj-04-2017-0057-Figure15-1.png",
        "caption": "Figure 15 Containment boundary for peripheral milling",
        "texts": [
            " For peripheral milling, the projection of the planar surface to the cutting plane will degenerate to a zero-area line. In this case, a line fitting through the projected vertices is first calculated, from which a rectangular tool containment boundary will be generated. We assume that the fitted line is defined by two points Pa, Pb, the normal of this planar surface is N, the machining allowance on this surface is t and the diameter of the selected tool is D. The four points of the rectangular tool containment boundary is shown in Figure 15 and calculated as the follows: Pa 0 \u00bc Pa 1N t1 D 2 1D Pb 0 \u00bc Pb 1N t1 D 2 1D Note that, theoretically, D should be zero. In practice D is set to a constant small value obtained from trial and error (0.01 inches in practice) to compensate for any inaccuracy of the data in this process. The influence of the value of D on the machining process for CNC-RPHybrid is trivial, as it simply ensures adequate access of the tool to the surface, even in the presence of small errors in tessellation. Similar to planar features, tool path planning for cylindrical holes depends on the parametric information that is extracted from the color model approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003949_s00773-019-00660-1-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003949_s00773-019-00660-1-Figure6-1.png",
        "caption": "Fig. 6 Internal frame design of UTP-AUG",
        "texts": [
            " (31) output = [ rpx \u2212 rpxd rpz \u2212 rpzd mb \u2212 mbd ]T . (32) mbd = (m \u2212 mh \u2212 m\u0304) + 1 g (\u2212 sin \ud835\udefed(KDo + KD\ud835\udefc 2) + cos \ud835\udf09d(KLo + KL\ud835\udefcd)V 2 d , (33) rpxd = \u2212rpzd tan d + 1 mg cos d ((mf3 \u2212 mf1)vxdvxd + (KM0 + KM d)V 2 d ), m1d = mbd + mh + mf1, mod = mbd + mh + m\u0304 \u2212 m, m3d = mbd + mh + mf3, (34) d = 1 2 KL KD tan( d) \u239b\u239c\u239c\u239d \u22121 + \ufffd\ufffd\ufffd\ufffd 1 \u2212 4 KD K2 L cot( d) (KDo cot( d) + KLo) \u239e\u239f\u239f\u23a0 . 1 3 A prototype of UTP-AUG, with fixed wings, rudder, internal ballast actuator, on board electronics and batteries has been developed, as shown in Fig.\u00a06 with the parameters shown in Table\u00a01. The wings are fixed to the vehicle, however wings of different sizes can be attached to investigate the vehicle\u2019s lift to drag ratios [16]. The pitch of the glider is controlled by an internal moving sliding mass along a linear guide way of 20\u00a0cm in length and 825\u00a0cm3 Dive system Tmax ballast tank with 12\u00a0V DC motor driver. The ballast tank and linear actuator are fitted in the frame, as shown in Fig.\u00a06. DC servomotor controls the position of internal moving battery mass by using limit switches. The polarity of the DC servo is changed when the desired diving depth is reached. The relative positions of the ballast tank and internal moving mass allows the moderate adjustments in the position of glider\u2019s center of gravity (CG). The high torque servos are driving the plungers by connecting it to linear potentiometers. A Closed loop feedback controller is used to regulate the tank\u2019s plunger position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure13-1.png",
        "caption": "Figure 13: z-displacement of the flux of force adapted lattice structure with straight struts",
        "texts": [
            " Structure Geometry. To obtain the described geometry with straight struts, the nodal points of the previous structure have been retained and the curved structure has been replaced by straight beam elements. Analogue to the investigations presented before, the diameters of the struts were set to 2 mm. The resulting structure and its constraints and loads can be seen in figure 12. The mass of this geometry is 83.58 g. Thus, the weight is similar to the curved structure and 43 % below the periodic one. Stiffness. Figure 13 shows the simulation result for the z-displacement of the structure. In order to make the results more comparable, the colour scale is the same as in figure 5 and 9. The displacement in the point of force application is 0.130 mm. Together with the applied force of 300 N, a stiffness of 2317 N/mm results for the structure. This means an improvement of 7 % compared to the curved structure and 9 % compared to the periodic geometry. This leads to a value of 27.7 N/(mm*g) for the stiffness to mass ratio"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001585_aim.2018.8452392-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001585_aim.2018.8452392-Figure7-1.png",
        "caption": "Figure 7. Joint part in the previous robot",
        "texts": [
            " In addition, the retaining ring is fastened with a nut to fix the artificial muscle. By using this method, the binding force against the expansion of the artificial muscle can be enhanced, and slip-off of the artificial muscle can be prevented. The robot also includes a joint portion. This joint has a structure imitating a universal mechanism. For this reason, the Amount of contraction 56 mm 54 mm 118 mm 978-1-5386-1854-7/18/$31.00 \u00a92018 IEEE 937 degree of freedom of the joint is high and it is possible to pass through the bent pipe (Fig.7). A spring is contained inside to increase its rigidity. In addition, solenoid valves are stored inside the joint. The robot was also problematic in that the amount of contraction in the unit part alone was small, and the air pressure response was poor. The robot was difficult to handle when put to practical use. A peristaltic motion-type robot was developed with the aim of improving the motion of the robot with a view toward practical applications and improving basic performance. The authors developed another peristaltic motion robot named PEW-RO V"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003767_jctb.5931-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003767_jctb.5931-Figure1-1.png",
        "caption": "Figure 1. Configuration of Microbial Fuel Cell(MFC) and reactions in anode and cathode.",
        "texts": [
            " The vitamin solution had the following composition: 2 mg biotin, 2 mg folic acid, 10 mg pyridoxine HCl, 5 mg thiamine HCl, 5 mg riboflavin, 5 mg nicotinic acid, 5 mg D-ca-pantothenate, 0.1 mg vitamin B12, 5 mg p-amionobezoic acid, and 5 mg lipoic acid, per liter DW. The MFCs were constructed from two media bottles (310 ml capacity; Corning, N.Y.) joined by a glass tube containing a 1.25-cm-diameter proton exchange membrane (PEM) (Membrane International Inc., USA, AMI-7001). A clamp was used to combine the two bottles(See Figure 1.). The anode electrodes were made of plain porous carbon paper (2 cm \u2179 5 cm, 10 cm2 area, NARA Cell-Tech, Korea) and the cathode electrodes consisted of 10% Pt-coated plain porous carbon paper (2 cm x 5 cm, 10 cm2 area, NARA Cell-Tech, Korea). The electrodes, membrane, and each of the chambers were washed using 5 N HCl and 5 N NaOH and, then, were rinsed with distilled water for impurity removal and sterilization. The growth medium (250 ml) was filled into the anode chamber and purged with 100% nitrogen for 10 minutes before inoculation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001956_citcon.2015.7122604-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001956_citcon.2015.7122604-Figure6-1.png",
        "caption": "Fig. 6: High frequency switching results in capacitive discharge across the bearings",
        "texts": [
            " The lubricant film acts as a dielectric charged from the rotor. For high frequency currents, it forms a capacitor. The capacitance depends on the type of lubricant - thickener and oil type - lubricant film thickness, temperature, and viscosity of the lubricant. When the 978-1-4799-5580-0/15/$31.00 \u00a9 IEEE 2014-CIC-1045 voltage reaches the voltage breakdown limit of the lubricant, the capacitor will be discharged and a high frequency current occurs. This discharge will occur every time the motor switches. Fig.6. Electrical discharge occurs through the lubricant separating the rings and rolling elements when current passes through the bearing. Heat is generated from the dischargeand local melting of the surface metal occurs. Craters or pits are then formed and local particles of molten material are transferred and partially break loose. The metal at the crater is rehardened and more brittle than the original surface metal. Below the rehardened material is a layer of annealed material which is softer than the surrounding material"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003270_s12008-017-0437-5-Figure16-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003270_s12008-017-0437-5-Figure16-1.png",
        "caption": "Fig. 16 Comparison of deformations between the proposed ChainMail algorithm (top row) and traditional ChainMail algorithm (bottom row): a cube extension; b cube compression; c cylinder extension; and d cylinder compression",
        "texts": [
            " 15 using shear deformation modelled by the proposedmethod and traditional ChainMail algorithm. It can be seen that the traditional ChainMail produced unrealistic deformation behaviours due to not conserving linear and angular momentums. Further, it can also be seen from Fig. 15 that the deformed shape produced by the proposedmethod behaves nonlinearly, whereas the traditional ChainMail behaves only linearly. The nonlinear deformation is attributed to the use of nonlinear bounding regions of ellipsoid shape and associated posi- tion adjustments. Additional comparisons are presented in Fig. 16. The proposed ChainMail algorithm exhibits significantly more volumetric behaviours than those of traditional ChainMail algorithm. With the proposed method, Fig. 16a shows that the side faces of the cubic model shrink inwards due to tensile deformation on the top face, and Fig. 16c shows that the bottom side of the cylinder model follows the movement of tensile deformation on the top side.With above deformation examples, all of the deformations exceed 10%of the original mesh size. This demonstrates that the proposed method can accommodate large deformation of soft tissues. Trials are also conducted to compare the computational time of proposed method with traditional ChainMail algorithm under same conditions. The computational time is evaluated on an Intel(R) Core(TM) i5-2500k CPU@4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000025_1.3667502-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000025_1.3667502-Figure7-1.png",
        "caption": "Fig. 7 Model details and thermocouple locations",
        "texts": [],
        "surrounding_texts": [
            "HALF PLANE EDGE\no I\n/ HOLE BOUNDARY CIRCLE\nT T(y)= Tag(y) T(r)= Tof(r)\na) HALF PLANE AND HOLE\nured into the plate f rom the boundary, either normal to t he straight edge or radially in t he fillet region.\nAnalysis. A comparison of the stresses parallel to the straight and curved boundaries may be obtained from the rat io of equations (7) and (21),\no'e/o'x = f i f ) - (R + f)~ ! J > + p)J\\p)dp\ng(f) (22)\nIn the limiting case of a perfect thermal shock with R > 0, f = 0, and fir) = g ( f ) = 1, as a result of which cTe/o-x \u2014 1.\nIn the general case the integral will not vanish as i t did for perfect thermal shock and fur thermore f ( f ) y^ g(r). I t will now be shown t h a t when a proper limitation is placed upon the penetration distance s of the temperature field little chauge occurs in the character of the stress field in the region of the corner fillet compared to t h a t in a half plane.\nTor an axisymmetric temperature field\nb2T/df2 + (R + f)~\\DT/b?) = (1 /D)(dT/dd) (23)\nwhile for a uniaxial Cartesian field,\nb2T/df2 = (l/D)i>T/bd (24)\nFor the purpose of this demonstration assume t h a t a t a given 9,\nT = T0e-cr (25)\nwhere c > 1 to insure rapid decay of T in the f-direction. Some justification for the choice m a y be found in an article by Heisler [5]. Then referring to Fig. 4\n{R + ?)-'(c)27c>F) = (R + f)~\\bT/by) =\n-cT0e-\u00b0\\R + f)~l (26)\nd2T/bf2 = (<)2T/by2) = c2Toe\"\" (27)\nand therefore the ratio of the two terms, defined as (3, is\n(R + f)-'(ar/Sr) d'T/dr2 c(R + f)\n(28)\nIn order to set a mathematical limit to the penetrat ion distance, assume t h a t cr = cs = 4 so t h a t T/To = c~4 a t s = mR where m is a number less than 1. Then c = 4 / m R and\njS = mR\n4 (R + mR) 4(1 + m) (29)\nThen it follows t h a t for m in the order of l / t to l/i (or s ^ (R/2)) the term (R + f ) ~l(dT/dr) would be significantly less than Z>2T/dr2 and the temperature fields for the infinite half plane and the fillet would not differ significantly under thermal shock. For this case, then, the f-wise stress field in the half plane would be virtually the same as tha t in the perforated infinite plane, and consequently the union of a segment of each structure to form a filleted corner should result in a continuous stress field which is essentially constant along the entire contour of the space so formed. The only limitation to be applied a t present is t ha t the shock penetration depth s should be less than half the fillet radius.\nTo estimate the difference between the tangential and plane stresses, assume t h a t fir) = g(r) = c~cr. Then equation (22) becomes\nag_ fx = 1 - (R + ry\nWhen integrated, this becomes\nf Jo (R + p)e~ cpdp\nvo_ = i _ je\" - 1 )(cR + 1) o-x (R + f)*c2\n(30)\n(31)\nFor R = '/4 in., cs = 4, and s = R/4, the stress ratio would vary with r as shown in Fig. 5, from which it is evident t h a t the stress field a t the fillet corresponds close!}' to the half-plane stress field in the region near the boundary.\nFor a half-plane penetration depth s equal to R or greater, or in the case of a very small fillet radius, t he penetration in the punctured space would be much less than t h a t in the half plane, and consequently t he inward extent of the curved stress field would be. much less than the straight stress field. Furthermore, the boundary temperature in the curved portion would lag behind t ha t in the straight portion. As a consequence, the preceding analysis would not apply.\nExperimental Results. I t has been indicated t ha t if the shock penetration depth s were less than half the fillet radius the stress field would be essentially linear. If the shock penetrat ion depth is defined as the distance s into the space a t which T = 2'oe-4 or roughly 0.027'fi, it affords a slightly more precise definition of the penetration distance than provided previously.\nMeasurements of this distance on the straight edges of the Homalite 100 corner model yielded a value of s = 0.12 in. a t 6 = 5 seconds. Consequently, it was anticipated t h a t for R = l / 4 in. no significant change in the character of the thermal shock stress\n3 4 6 / A U G U S T 1 9 6 2 Transactions of the A S M E\nDownloaded From: http://manufacturingscience.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmsefk/27464/ on 06/12/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
            "0 - SEC\n2\nFig. 6 Notch fringe patterns at indicated times\nfield would be observed along the model boundary including the fillet region. Tha t is, the radial variation of fringe order would be the same as observed normal to a half-plane boundary. This is substantiated in Fig. 6, which depicts the fillet fringe patterns for several times. The boundary-layer effect is clearly evident.\nThe transition in the stress distribution from a half-plane field to a three-quarter plane field is evident in Fig. 6, which reveals the variation in fringe pattern in the region of the corner. This result is in accordance with theory. Furthermore, Fig. 6 reveals the beginning of encroachment of the stress field into the interior of the 3/i space near the fillet as 6 became large (15 sec, 30 sec).\nThree-Dimensional Effects. As long as the thickness of an experimental model is finite, no matter how small, the theoretical possibility of a three-dimensional stress field in the region of the corner must be admitted. Such an effect would be confined to the region of the fillet. However, the radius would need to shrink toward zero before this effect would become significant.\nConclusions General Conclusion. As n result of the theory and corresponding experimental verification, evidence has been presented that aETo is the upper bound to the stress induced in an arbitrary elastic, simply connected plane space by a thermal shock temperature change To if the penetration distance s of the temperature field is less than half the local radius of curvature of the space boundary. Under these conditions the stress field is essentially confined to a boundary layer of thickness s independent of the boundary shape.\nCorollaries. The corollaries to this result for such a space and for s S R/2 are indicated by data discussed in the thesis:\n1 Boundary shape would not play a role in thermal shock stress analysis.\n2 The peak concentration factor for thermal shock is unity, which is obtained on a straight boundaiy.\nSummary. In view of these residts, the general validity of\nequation (2) appears in doubt since K is unity for the conditions previously described. I t is usually considerably greater than unity for mechanical loads.\nA P P E N D I X A Experimental Program\nIntroduction. The experimental program involved the application of a thermal shock to a photoelastic model and the recording of temperature and fringe pattern data 2 to 10 sec after shock application. This is in contrast to previous investigations (by Gerard and Gilbert [6], for example) in which the first data sampling occurred 30 sec after shock.\nThis section is devoted to a description of the experimental investigation and a summarj>- of thermal and mechanical properties of the models.\nDetails of Model Fabrication. During the exploratory study conducted bjr Gerard and Gilbert it was demonstrated that a relatively small photoelastic model can reveal the character of a thermal shock. In this program rectangular plates approximately 1 / i in. thick were used for studying thermal shock on an infinite half plane. Dry ice was applied to one plate edge while the faces were insulated by 1-in-thick styrofoam plates held against the model faces with double sided masking tape. The other model edges were left free. As Gerard and Gilbert demonstrated, almost the entire width of the plate was subjected to a uniform stress field. Consequently, for the purposes of the present study no preparation of the uncooled model edges was required.\nReduction of Infinite Half-Plane Data. T h e m e a s u r e m e n t s on t h e photographic negatives of the model fringe patterns were converted to curves of stress as a function of distance from the cooled edge for matching with the temperature curves. This was done by use of the stress optic law,\nJournal of Engineering for Industry A U G U S T 1 96 2 / 3 4 7\nDownloaded From: http://manufacturingscience.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmsefk/27464/ on 06/12/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
            "u ~ \"fit\n(32)\nThe temperature data were converted to stress by use of the relation\n(1 = aET (33)\nin which the product aE is obtained from calibration of a rectan gular strip in the manner described in [7).\nFinally, both Eels of data were plotted together IlS !1 function of position for each selected time. The discussioll of the results appears in the pertinent sections in the main body of the thesis.\nModel Materia l Properti es. The photoebstic model material prop~ erties are shown in Table 1.\nThe data on conductivity and diffllsivity were obta ined from Tramposch and Gerard [8L since they found these properties to vary relatively little.\nThe determination of the heat-transfer coefficient was made utilizing the chart of Fig. 8 which relates temperature, t.ime, and position for n. cOllvective heat-transfcr thermnl shoek on a half space tiS functions of thermal properties of the material in the ha lf space. By plotting 1'/1'u:lt the three thermocouple locat.ions for fJ = 10 scc, and interpolating to find (h/k)(DfJ) 112, t.he value. of h/k was found, from which h W[lS computed.\nExperimenta l Errors. The principal sources of experimental error were discrepancies in fringe order; inaccuracies in reading tem ~\nperatures duc to thermocouple location and calibration errors and signal indication; and variations in model material properties.\nThe fringe orders were obtained directly from the models . No attempt was made to extrapolate data to the edgE'S. Consc~ quently, fringe order errors would be limit.ed to inaccuracies in measurements of location on the model. This is estimated to be lE'sS than 2 per cent of the penetration distance, or about 0.002 in .\nThe largest source of error was in determining the location of\n348 / A U G U S T 1 962\nthe thermocouple lead. The vnriation in position could have been as large as 0.020 in. The possible error due to calibration and signal indication WliS 3 deg F . These effects arc evident in Fig. 9.\nCompa rison of th eory and\nMaterial property variations are roughly 10 pflr cent of the values sho<ivn in Ta.ble 1 for D and k, while aE/f would be ex pected to agree within 3 per cent of the values shown in Table 1.\nReferences H. Becker and H. Tramposch, \"Photothermoelastic Investiga tion of Notched Plat.es,\" New York University Report. 59\u00b71, January, 1959.\n2 W. Jaunzemis find E . Sternberg, \"Trnnsient Thermal Stresses in a Semi~Infinite Slab,\" Journal of Applied Meeham'es, vol. 27, THANS. ASi\\'IE, vol. 82, 1960, p . 93.\n3 C. K. Youngdahl and E. Sternberg. \"Transient Thermal Stresses in II. Circular Cylinder,\" ASlVIE Paper No. 60-WA~1 23, presented November, 19130.\n4 S. P. Timoshenko, \"Theory of Elasticity,\" McGfllw~Hill Book Company, Inc .. New York, N. Y., 1934.\n5 i\\L P. Heisler. \"Temperature Charts for Induction and Con~ stant Temperature Heating,\" TUANs. ASi\\'IE, vol. 69, 1947, pp. 227- 236.\nB G. Gerard and A. C. Gilbert, \"PhotothcrmoelusLicity: An Ex\u00b7 plorutory Study,\" JOIlT1wl of Applied Ilfechanics, vol. 24, Tn .... Ns. ASME. vol. 29,1957, pp. 3M-3GO.\n7 H. Tramposch, \"Theoretical and Experimental Investigation of Thermal Stresses in Hypersonic Aircraft 'Ving Structures,\" New York University Doctoral Thesis, New York, N. Y.\n8 H. Tramposch and G. Gerard, \"Physical Properties of Plastics for Photothermoelastic Investigations,\" JOllrnal of Applied AI eella/.ies, vol. 25, TRANS. ASi\\1E, vol. 80, 1958, pp. 525-528.\nTransactions of the ASME\nDownloaded From: http://manufacturingscience.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmsefk/27464/ on 06/12/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use"
        ]
    },
    {
        "image_filename": "designv11_13_0000413_s12206-011-1201-6-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000413_s12206-011-1201-6-Figure4-1.png",
        "caption": "Fig. 4. Deformation of the tapered roller bearing under bending moment.",
        "texts": [
            " According to lumped mass method, the beam is simplified to be several disks and beams, and the disk is considered to have no thickness and the beam is considered to have no mass. As Fig. 3 shows, the air blower rotor is supported by two tapered roller bearings. Under the action of radial load, the shaft bends, and bending moments by an external radial load will act upon the two bearings. Under the radial load, the shaft will produce radial elastic displacement and produce a tilting angle between the inner and outer raceways of bearings. As Fig. 4 shows, for the tapered rolling bearing, the contact angle is \u03b1 , conical angle is 2\u03b2 and the angle between neighbor roller center lines is \u03c8 . Under combined actions of radial load and bending moment, the bearing will deform. Each rolling element is divided into n sections uniformly along the generatrix, and then they are analyzed separately according to the upper and lower half circle. ( )1 2 L Li i n \u03b4 \u03b8 \u239b \u239e= \u2212\u239c \u239f \u239d \u23a0 (1) where 0, 2 Lx \u23a1 \u23a4\u2208 \u23a2 \u23a5\u23a3 \u23a6 \uff0c 0q = when 0\u03b4 \u2264 . For the bottom roller, the deformation on the i th section is ( )2 r Li i n \u03b4 \u03b4 \u03b8= + \uff0c [ ]0,x L\u2208 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003093_s1068366617030163-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003093_s1068366617030163-Figure10-1.png",
        "caption": "Fig. 10. Results of analysis of hydrodynamics for a connecting-rod bearing of a UNP55-250 pumping unit without consideration of radial compliance: (a) trajectory of shaft motion; (b) pressure distribution with marked two-dimensional graph in the zone of maximum pressure.",
        "texts": [
            " The finite-element analysis of the lubricating f luid f low in connecting-rod bearings of a UNP55-250 pumping unit was performed takes into consideration the radial compliance of the shaft and bush. The bearing width is 70 mm, the diameter of the shaft is 92.21 mm, the diameter of the bush is 92.25 mm, the dynamic viscosity of the lubricating material is 0.0277 Pa\u22c5s, the radius of the crank is 34 mm, the length of a connecting rod is 245 mm, and the angular speed is \u03c9 = 43.03 rad/s. For the comparison, Fig. 10 presents the main results of the hydrodynamics analysis of an undeformable bearing; Fig. 10a shows the trajectory of the motion of the shaft and Fig. 10b shows the pressure profile for the position of the shaft when the maximum pressure value is realized. In all considered cases, this point corresponds to a 95\u00b0 angle of rotation CONSIDERATION OF THE MULTIZONE HYDRODYNAMIC FRICTION 249 of the shaft. For this case, the maximum pressure of 16.2 MPa and the minimum gap of 13.9 \u03bcm was obtained. In the connecting rod, the big end of the UNP 55- 250 plant, a commercial bush for tractor diesels with a width of 70 mm and an outer diameter of 100 mm is used",
            " The hydrodynamic calculations were performed using one-layer (antifriction layer) and two-layer (the antifriction layer and the base) compliant bushes. The compliance of both the solid and hollow shaft was also considered. Because the diameter of the lubricant supply hole in the bearing is small compared to the rest of the radial sizes of the parts, the influence of this hole on the hydrodynamic parameters is insignificant. The pressure profiles without taking into consideration the compliance are given in Fig. 10. The consideration of the compliance of a bearing with an antifriction layer made of AO20-1 alloy leads to a decrease in the maximum pressures to 10%. In the case of using metal\u2013polymer bushes, the compliance factor is more significant. In Table 4, the results of the hydrodynamic calculations of a bearing with an antifriction layer made of polymeric composite in accordance with the CRR scheme (a compliant ring in a rigid holder) are presented. To provide the visual graphic presentation of the calculation results with a compliant antifriction layer, the new interface of the second version of the BBFEM software was developed in which the pressures are presented in the form of the dimensional profile (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000157_socpar.2014.7007976-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000157_socpar.2014.7007976-Figure1-1.png",
        "caption": "Fig. 1. REB geometry.with damaged outer race.",
        "texts": [
            " The contents of this paper are organised in the following way. In Section 2, background on REB signals is introduced. In Section 3, both the methodology proposed to detect and diagnose faults in bearings and the description of the techniques involved in the main steps of the method are presented. In Section 4, the proposed method is validated using experimental bearing data. Finally, the conclusion of this work are given in Section 5. II. NONSTATIONARY NATURE OF DEFECTIVE REBS VIBRATION RESPONSE As shown in Fig. 1, a REB has four major components, outer race, inner racer, rolling elements, and cage. All four components might be damaged during operation. Generally, the signature of a damaged bearing consists of exponentially decaying ringing that occurs periodically at the characteristic frequency. The vibration signal of a defective bearing usually considers being amplitude modulated at characteristic defect frequency [3]-[6]. A large number of models have been used to describe the dynamic behavior of rolling element bearings under different types of defects",
            " Shock-pulses are generated each time rolling elements strike the defected surface of the bearing outer race and consequently, excite resonances of the structure between the fault location and the vibration sensor [3]. The frequency of shock occurrence is usually called the ball pass frequency of the outer race (BPFO) and is expressed as: 1 cos 2 br c DNf BPFO D \u03b8= \u2212 , (1) where fr is the shaft speed (in Hz), N is the number of rolling elements, is the angle of the load from the radial plane. Additionally, Dc denotes diameter between two opposite rolling elements and Db denotes the diameter of a single rolling element (see Fig. 1). Assuming constant rotational speed and load of the bearing, the vibration signals generated by a REB with defected outer race were modeled by Randall et al. as [3], [6]: ( ) ( )1 i i x t t i n t BPFO \u03c9 \u03c4= \u2212 \u2212 + , (2) where (t) is the waveform generated by a single impact (related to resonance frequencies of the system), i is an independent and identically distributed random variable and n(t) is an additive background random noise. It now has to be stated that i introduces influence of the rolling elements slips into the model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001992_ecce.2015.7310061-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001992_ecce.2015.7310061-Figure1-1.png",
        "caption": "Fig. 1: IPM Motor 1 and 2 considered for analysis",
        "texts": [
            " In [8], a 2D analytical model developed by the authors which was extended to 3D using an end-effect factor. An analytical model in 3D for estimation of magnet eddy current losses in IPM machine is still missing. In this paper, analytical and numerical finite element analyses of eddy currents in the magnets of IPM machines are presented. The analyses consider the effect of segmentation in the axial and tangential directions. The analytical model can be used in choosing the appropriate magnet segmentation numbers to reduce the eddy current losses in an IPM motor. II. SPECIFICATIONS OF IPM MOTORS Fig. 1 shows the two types of IPM motors considered for analysis [17]. Both motors have 8 poles and an internal and external stator diameter of 161.93 mm and 269 mm respectively. The air-gap length and axial length were 0.73 mm and 84 mm respectively. The magnet in both motors has a thickness of 6.5 mm while the width was 37.8 mm for Motor 1 and 18.9 mm for Motor 2. The Cartesian co-ordinate system is chosen over the cylindrical co-ordinate system for the analysis as the magnets in an IPM machine are flat shaped, thereby reducing the complexity of the analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002392_j.jcp.2014.01.008-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002392_j.jcp.2014.01.008-Figure1-1.png",
        "caption": "Fig. 1. Dual IB representation of an immersed surface.",
        "texts": [
            " The second component of this approach is to introduce penalty forces that act on the corresponding points of the two immersed boundaries to oppose motions that move them apart. The details of our two-fluid IB method follow. As in the standard IB method, we use an Eulerian description of the fluid variables, in this case for each of the two fluids, and we use a Lagrangian description for each of the immersed boundaries. For each immersed structure, we introduce two immersed boundaries, which we denote \u0393 n and \u0393 s, as shown in Fig. 1. The locations of points on the solvent and network immersed boundaries are given, respectively, by the vector functions Xn(q, t) and Xs(q, t), with Xn(q,0) = Xs(q,0). Each Lagrangian point on \u0393 n communicates only with the network and each Lagrangian point on \u0393 s communicates only with the solvent. In particular, the velocity of an IB point on \u0393 n is determined by the local network velocity field un, while the velocity of an IB point on \u0393 s is determined by the local solvent velocity us. These velocities are specified by formulas similar to Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002523_j.measurement.2016.02.017-Figure16-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002523_j.measurement.2016.02.017-Figure16-1.png",
        "caption": "Figure 16 Torques and forces effected on Z-axis linear guideways",
        "texts": [
            " Then: The free body diagram shown in Figure 14. The total friction force in X-axis is as following: Eq. 6 Where; (refer to Figure 8). The static and kinetic coefficient friction ( ) are considered to identify the friction force in X- axis in both of stationary and moving table on the guideways. Eq. 7 Eq. 8 Where, is a total static friction force and is total kinetic friction forces in X-axis linear guideways. The cutting force and normal force result in the friction force on Z direction during cutting operations are shown in Figure 15. Figure 16 shows the force and torque affected on Z-axis direction. It is necessary to transfer the moment of the forces due to cutting forces and normal force, to the indevidual center of gravity in order to calculate the forces acting on the linear guideways. From Figure 15 to Figure 17 (a), Eq. 9a Eq. 9b From Figure 15 to Figure 17 (b), Eq. 9c Eq. 9d From Figure 15 to Figure 17 (b), Eq. 9e Eq. 9f From Figure 15 to Figure 17 (c), Eq. 9g Eq. 9h Free body diagram for Z-Axis linear guide after analysis the forces is shown in Figure 18"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001786_9783527344758-Figure7.7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001786_9783527344758-Figure7.7-1.png",
        "caption": "Figure 7.7 Added mass coefficient for symmetric geometries.",
        "texts": [
            "5) Microtubules: Nanobiomechanical Simulation 185 F = \u23a1 \u23a2\u23a2\u23a2\u23a3 m11 m12 m16 m21 m22 m26 m61 m62 m66 \u23a4 \u23a5\u23a5\u23a5\u23a6 \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239d du1 dt du2 dt du6 dt \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u23a0 (7.6) F = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 m11 m12 m13 m14 m15 m16 m21 m22 m23 m24 m25 m26 m31 m32 m33 m34 m35 m36 m41 m42 m43 m44 m45 m46 m51 m52 m53 m54 m55 m56 m61 m62 m63 m64 m65 m66 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239d \u2022 u1 \u2022 u2 \u2022 u3 \u2022 u4 \u2022 u5 \u2022 u6 \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u23a0 (7.7) For symmetric geometries, this tensor is extremely simplified. For example, for the circle, the ellipse, and the square, these values are shown in Figure 7.7. For a three-dimensional sphere, the values are as follows, and the remaining mass terms will be zero: m11 = m22 = m33 = 1 2 \ud835\udf0c\u2200 = ma (7.8) According to the above description, since the microtubule is submerged in cytoplasm, the reaction of the fluid around it should be considered. Due to the fact that tubulin is considered as a sphere, the added mass will be added to the mass of tubulin for the surrounding fluid. According to above equation, the added mass will be equal to 1 2 \ud835\udf0c\u2200 where \ud835\udf0c is density of cytoplasm and \u2200 is volume of tubulin"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003435_1.5034589-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003435_1.5034589-Figure1-1.png",
        "caption": "FIGURE 1. The cylinder and the coordinates.",
        "texts": [
            " The system of equations (1) for finding the reactions of walking robot body and legs can then be transformed [7, 8, 9] to\u2211n i=1 R\u0303i = \u03a6\u0303, \u2211n i=1 r\u0303i \u00d7 R\u0303i = r\u0303C \u00d7 \u03a6\u0303. (2) We now assume that robot footholds are on a rough cylinder of radius \u03c1, with a friction coefficient k. Let \u03c1 be the cylinder radii. Introducing (right-handed) coordinate system O\u03be\u03b7\u03b6 so that an axis O\u03b6 be parallel to the force vector \u03a6\u0303, and the coordinate system Oxyz with Ox along the cylinder axis, so that \u03a6\u0303 have a negative projection on axis Ox \u2013 see Fig. 1. Here, the axis Oz is directed so that the plane Oxz is parallel to the vector \u03a6\u0303. Then the cylinder axis has an angle \u03b1 with vector \u03a6\u0303, and the planes O\u03be\u03b6 and Oxz are identical. The task of solving Eqns. (2) is similar to the reactions distribution footholds problem in the homogeneous gravity field, when the footholds are on the external surface of a rough inclined cylinder whose axis has an angle \u03b1 with the vertical vector \u03a6\u0303, and is related to Ref. [7], where the problem of calculating the reaction components along the cylinder axis when \u03b1 = 0 and the work [9] for arbitrary \u03b1 was considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003460_s12206-018-0441-0-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003460_s12206-018-0441-0-Figure4-1.png",
        "caption": "Fig. 4. Subproblems of the Paden\u2013Kahan definition of inverse kinematics: (a) Rotation about a single axis; (b) rotation about two subsequent axes; (c) rotation to a given distance [10].",
        "texts": [
            " Paden-Kahan subproblem The Paden-Kahan subproblem is the geometrical method used to obtain a rotation angle(\u03b8) about an axis as subproblems use the formula of the product of exponential (POE) based on the screw theory. The POE formula is a forward kinematic solution, which was introduced by Brockett [17]. The detailed explanation on the Paden\u2013Kahan subproblem was introduced by Murray et al. [10]. The classical subproblem is explained based on three motion types. The first is rotation about a single axis. The p is the point before the rotation, and the q is the point rotated by \u03b8 with respect to the axis vector S. The trace of q draws a circle, such as that shown in Fig. 4(a). The axis vector means the direction of rotation, and magnitude of axis vector are 1. The rotation angle(\u03b8) which is the solution of inverse kinematics is expressed as follows [10]: \u03b8 atan2( ( ), ) p p p p u v u v= \u22c5 \u00d7 \u22c5S (10) where up and vp represent the projection of u and v onto the plane normal to the axis vector. We used the arctangent2 function because that is covered the range of -180\u00b0 to 180\u00b0 The second is rotation about two subsequent axes. Point p is rotated about a second axis vector S2 until the point coincides with point pa",
            " In this case, to solve the inverse kinematics, pa is calculated by ( )2121 SSSS \u00d7++= \u03b3\u03b2\u03b1ap (11) where the magnitude of the vectors on each axis is unity, and ,\u03b1 \u03b2 and \u03b3 are given by ( )( ) ( ) 1 2 2 1 2 1 2 1 u v \u03b1 \u22c5 \u22c5 \u22c5 = \u22c5 S S S S S S - - (12) ( )( ) ( ) 1 2 1 2 2 1 2 1 v u \u03b2 \u22c5 \u22c5 \u22c5 = \u22c5 S S S S S S - - (13) ( )2 2 2 1 2 2u\u03b3 \u03b1 \u03b2 \u03b1\u03b2= \u00b1 \u22c5S S- - - (14) from these results, the solution of inverse kinematics is obtained from 2 2\u03b8 S a e p p    = and 1 1\u03b8 S a e q p \u2212   = using subprob- lem 1. Subproblem 3 is rotation to a given distance (see Fig. 4(c)). Point p is rotated about S2 until point p corresponds with q \u2013 \u03b4. The distance \u03b4 is given, and the solution is 0 \u03b8 \u03b8 \u03c6= \u00b1 and expressed as follows: ( )( ) 2 2 2 \u03b8 tan 2 , cos 2 p p p p p p p p p u v a u v u v a u v \u03b4 + \u2212 = \u22c5 \u00d7 \u22c5 \u00b1     S (15) where up, vp and \u03b4p, are respectively the projections u, v and \u03b4 onto the plane normal to the axis vector. Ones again, Subproblem1 is a geometrical method to find the inverse kinematic solution for a single axis (Fig. 5(a)), and subproblem2 can finds the intersection point for two consecutive axes and finds the rotation angle using subproblem1 (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003655_j.ins.2018.10.062-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003655_j.ins.2018.10.062-Figure1-1.png",
        "caption": "Fig. 1. Two inverted pendulums systems connected by a moving spring mounted on two cars.",
        "texts": [
            " In fact, the first terms on the right hand of the inequalities (79) and (80) are needed to be com- pensated, and the controller gains are decreased at the expense of increasing the power of the second terms of (79) and (80) . However, by using the graph theory, the effects of the second terms can be eliminated by constructing the global Lyapunov functions (87) , and the details are shown in (89) . 4. Example To show the effectiveness of the proposed control methods, two examples are considered in this section. Example 1. Example 1 : Here, referring to [22] , two inverted pendulums systems connected by a moving spring mounted on two cars with stochastic disturbances are considered, which are shown in Fig. 1 . The input to each pendulum is the torque u i (i = 1 , 2) applied at the pivot point. Define the state vectors as (x 1 , 1 , x 1 , 2 ) T = (\u03b81 , \u02d9 \u03b81 ) T (rad, rad/s) and (x 2 , 1 , x 2 , 2 ) T = (\u03b82 , \u02d9 \u03b82 ) T (rad, rad/s). The dynamics equations of the two inverted pendulums subsystems are described as \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 dx 1 , 1 = x 1 , 2 dt dx 1 , 2 = (( g cl \u2212 ka (t)(a (t) \u2212 cl) cml 2 ) x 1 , 1 + u 1 cml 2 + ka (t)(a (t) \u2212 cl) cml 2 x 2 , 1 + \u03b2\u03041 x 2 1 , 2 + ka (t)(a (t) \u2212 cl) cml 2 ( sin (2 t) \u2212 ( sin (3 t) + L )) ) d t + cos (x 1 , 2 ) d w y 1 = x 1 , 1 \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 dx 2 , 1 = x 2 , 2 dt dx 2 , 2 = (( g cl \u2212 ka (t)(a (t) \u2212 cl) cml 2 ) x 2 , 1 + u 2 cml 2 + ka (t)(a (t) \u2212 cl) cml 2 x 1 , 1 + \u03b2\u03041 x 2 2 , 2 + ka (t)(a (t) \u2212 cl) cml 2 ( sin (3 t) + L \u2212 sin (2 t)) ) d t + sin (x 2 , 2 ) d w y 2 = x 2 , 1 where a (t) = sin (5 t) , c = m/ (m + M) , \u03b2\u03041 = m sin (x 1 , 1 ) /M, \u03b2\u03042 = m sin (x 1 , 2 ) /M"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000555_iecon.2012.6389491-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000555_iecon.2012.6389491-Figure4-1.png",
        "caption": "Fig. 4. Optical system.",
        "texts": [
            " It is worth noting that the response of commercial solutions for real-time control of robots is much slower than that of the laser system. As mentioned in Section I, although in exiting industrial systems the control of the cladding process is usually based on temperature measurements, research in the area is increasingly based on the use of CCD or CMOS cameras. In this context, FPGA-based systems have been proven [5][7] to be a better solution than existing industrial PC-based ones. The optical system (Fig. 4) is placed in a coaxial arrangement with the laser beam. It consists of a CMOS camera with 8-bit (256 grayscale levels) CameraLink interface and dynamic range up to 120dB, and auxiliary optical components: mirrors, a lateral augmentation telescopic system, and two filters, one to protect the camera from the laser beam and another to eliminate all components outside the visible spectrum. The FPGA system, whose block diagram is shown in Fig. 5, is in charge of the image processing and control algorithms, as well as of generating the corresponding actuation signals"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure3-1.png",
        "caption": "Figure 3: Design space, constraints and load for optimization",
        "texts": [
            " For this reason, new methods and algorithms have to be developed in order to adapt mesoscopic structures like honeycombs or lattice structures to the flux of force inside of a part. In the following sections, the benefit of such a strategy for the optimization of lattice structures will be investigated. In the following sections, lattice structures, which are adapted to the flux of force inside a part, are compared to a structure with a regular build-up. For that purpose, a design space with constraints and forces as depicted in figure 3 is used as initial situation for the respective optimizations. This design space is exposed with a fixed bearing on the one end and a shearing load of 300 N on the other end. Thereby, the Force is distributed on the respective surface by the use of ideal rigid RBE2 (Rigid Beam Element 2) elements. The examinations are performed by Finite-Element-Analyses (FEA) in order to calculate the mechanical properties and the potential for lightweight design of the particular structures. The material used in the simulation is aluminium 6061 without thermal treatment, which has the following mechanical properties: - Young\u2019s modulus: 58",
            " It is assumed that the maximum stress linearly depends on the applied force on the structure. Hence, it can be calculated under which load the limit of elasticity or the breaking limit are reached. Similar to the stiffness, the ratio of this load to the respective mass is used to compare the different structures, whereat a higher ratio indicates a better lightweight design, again. Furthermore, the stress conditions inside the structures are analysed by a sectional view through the cutting planes 1 and 2 (see figure 3). Plane 1 cuts the design space perpendicular to the y-axis in its middle. Plane 2 cuts the design space perpendicular to the x-axis, 60 mm away from its bearing. Initially, a periodic structure build-up is investigated. This is the currently established approach for the application of lattice structures in lightweight design. Several software tools for the generation of such geometries for additive layer manufacturing are already available. Structure Geometry. The structure is created by joining one and the same cubic elementary cell in the three directions in space. As mentioned before, a strut diameter of 2 mm and an interval of 5 mm have been chosen. The explained build-up has been applied on the design space in figure 3 (see figure 4). This leads to a total mass of 147.5 g for the whole structure. The bearing and the force application are realized as described before. Stiffness. Figure 5 shows the simulation result for the z-displacement of the structure in figure 4. The displacement in the point of force application, which is the reference point to compare the different structures, is 0.142 mm. Together with the applied force of 300 N, a stiffness of 2117 N/mm results for the structure. To make the performance more comparable to other structures, the ratio of stiffness to mass is calculated",
            " In figure 6 a), the von-Mises comparison stress is depicted for the regular lattice structure under the given load. The maximum value for the von-Mises stress appearing in the simulation is 50.58 N/mm\u00b2. Under the assumption, that this maximum stress linearly depends on the applied force, a maximum load of 406.6 N can be calculated for the obtaining of the limit of elasticity. Hence, the ratio of this force to the part\u2019s mass is 2.76 N/g. This value will be used in order to compare the structure with the further ones in the following sections. In figure 6 b), the sectional view at cutting plane 2 in figure 3 is depicted. Thereby, the stress in xx-direction for the cross section of six central struts in this plane is investigated. It can be recognized, that in the middle of the struts, only minor stresses appear. In contrast, in the upper sections pull forces and in the lower sections push forces can be observed, whereat both have similar absolute values. This leads to the conclusion, that there exists a major bending load on the horizontal struts. A similar behaviour can be observed in the outer struts of the specimen, even if the bending load is superposed with a push or pull force there, dependent on the position of the strut inside the structure",
            " This was explained with the help of the main stress tensors (respectively the flux of force) of a torsion-loaded hollow shaft, which also run along a helical path with an angle of 45\u00b0 compared to the shaft axis. Thereby, the appearance of bending loads within the single struts was avoided, which can be explained with the fact, that inside a loaded part, there do not appear any shear loads along the main stress directions [10]. This principle will be applied for the build-up of a flux of force adapted lattice structure in the design-space with its constraints and loads (see figure 3). Based on the results presented in [12], the lattice structure for the shear-loaded beam in figure 3 has been adapted to the flux of force inside the design space. For this, a MATLAB-based algorithm has been developed, which determines the courses of the flux of force along the first, second or third main stress direction from a given starting point. The input for this algorithm is the solver-output data of a Finite-Elements-Analysis of the design space under the defined constraints and loads. The procedure of the algorithm is based on an approach, presented in [10] for the optimization of fibre-reinforced plastics components. Structure Geometry. The course of the lattice structure inside the design space (see figure 3) has been adapted to the internal flux of force for the given constraints and loads (see figure 8). In order to keep the results comparable, the struts\u2019 diameters have been chosen to 2 mm as in case of the periodic structure. So, the course of the structure is the only differing variable in contrast to the periodic structure. The resulting structure has a mass of 83.57 g, which is 43 % less than in case of the periodic structure. Stiffness. Figure 9 shows the simulation result for the z-displacement of the structure",
            " Together with the applied force of 300 N, a stiffness of 2166 N/mm results for the structure. Due to the extensive reduction of the structure\u2019s mass, the ratio of stiffness to mass improves for 81 % towards a value of 25.9 N/(mm*g). Maximum Stress. In figure 10, the von-Mises comparison stress and the stress in xx-direction in cutting plane 2 are depicted for the flux of force adapted lattice structure. Figure 10 b) shows the stress in xx-direction for the cross section of the struts in cutting plane 2 (see figure 3). In order to make the results more comparable, the same colour scale has been used as in figure 6 b). It can be recognized, that only minor stress variations appear inside the single struts. This leads to the conclusion, that the bending loads in this sectional view have severely been reduced. Nevertheless, the maximum value for the von-Mises stress has risen to 74.63 N/mm\u00b2 (see figure 10 a). Under the assumption, that this maximum stress linearly depends on the applied force, a maximum load of 221.1 N can be calculated for the obtaining of the limit of elasticity. This means a worsening of 46 % compared to the periodic structure. The reason for the increased maximum stress can be seen in figure 11. The figure shows the sectional view at cutting plane 1 (see figure 3). It can be recognized, that there appears severe uparching, especially in longer struts with strong curvature. This uparching leads to high stresses along the struts and especially to stress peaks near the nodes. Although the ratio of the force at the limit of elasticity to the structure\u2019s mass (2.65 N/g) is just slightly below the value for the periodic structure (2.76 N/g) due to the reduced mass, this result is not satisfying. For this reason, a modified version of this structure with straight struts will be examined in the following section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003765_iros.2018.8594316-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003765_iros.2018.8594316-Figure7-1.png",
        "caption": "Fig. 7: Variable rigidity mechanism.",
        "texts": [
            " The pair of parallel springs are bended when oppositing and the third spring are bended when both oppositing and adducting. In the other words, when adducting only one spring is bended as the flexion of other general joints, but when oppositing two serial pairs of springs are bended. In this way the joint with wide range of motion is realized without little affection to other joints. D. Variable Rigidity Mechanism by Agonism Tendon Drive The MP joints of four fingers have variable joint stiffness mechanism as shown in Fig. 7. The wire from an actuator branches to palmar interossei muscle and dorsal interossei muscle. The rigidity of the rotational DOF of abduction and adduction is changed by tightening these two muscles. This mechanism is inserted in four fingers, but the only one actuator drives these muscles because the wires of four fingers are connected on the back of the hand. This mechanism cannot move MP joints actively but it is expect for fingers to bend along objects when the wire is loosen to make the rigidity of MP joints low and to transmit a force with the side of finger when the wire is tighten to make the rigidity high"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000203_j.ijmecsci.2013.11.010-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000203_j.ijmecsci.2013.11.010-Figure6-1.png",
        "caption": "Fig. 6. Schematics of braiding pattern for discussing rotational mechanism of threads.",
        "texts": [
            " Here, the length for each side of the rhombus can be expressed in terms of l0 as follows: AD \u00bc jDBj \u00bc jBCj \u00bc jCAj \u00bc b 2mn \u00bc l0: \u00f05\u00de Based on the fact that the bending deformation of a thread is much larger than the axial deformation, it is assumed that the length l0 remains unchanged during the actuator's deformation. In the lengthening and shortening deformation behavior of a McKibbentype actuator, the angle \u03b3 changes as the threads mutually rotate around each other during the deformation. Also, since the length of the side of the rhombus in the mesh remains unchanged, the rotational center does not move during the deformation. Next, the force T applied to a single thread in a sleeve as shown in Fig. 6 is considered. Here, the force T is decomposed into the components T\u03b8 and Tx in the directions of \u03b8 (circumferential direction) and x (axial direction), respectively. Also, the stresses in x- and \u03b8-directions acting on the rubber are termed sx;rub and s\u03b8;rub, respectively. The static equilibrium of forces in two directions are given as follows: T\u03b8\u00fes\u03b8;rub tRl0 cos \u03b3 \u00bc P\u2032Rkl0 cos \u03b3; Tx\u00fesx;rub tRl0 sin \u03b3 \u00bc \u00f0F\u00fe\u03c0R2 kP\u2032\u00de l0 sin \u03b3 2\u03c0R ; 8>< >: \u00f06\u00de where tR is the tube thickness as shown in Fig. 2(b), and Rk is the radius subtracting the thickness as Rk \u00bc R tk . Since the Young's modulus for the rubber is much smaller than that for the thread, the stress components sx;rub and s\u03b8;rub in Eq. (6) can be negligible. Therefore, the forces T\u03b8 and Tx can be simplified as follows: T\u03b8 \u00bc \u03c0RRkP\u2032 m tan \u03b3 ; Tx \u00bc \u00f0F\u00fe\u03c0R2 kP\u2032\u00de 2m : \u00f07\u00de Finally, the force T acting along a thread can be written as follows: T \u00bc T\u03b8 sin \u03b3\u00feTx cos \u03b3 \u00bc \u03c0R2P\u2032 cos \u03b3 m 1 2 Rk R 2 \u00fe Rk R \u00fe F 2\u03c0R2P\u2032 \" # : \u00f08\u00de As can be imagined from Fig. 6 that the shortening deformation of the actuator occurs due to the force T\u03b8 , which causes the threads to rotate and increases the angle \u03b3. However, the moment due to the force Tx works as a resistance. On the other hand, the lengthening deformation occurs due to the force Tx, which causes the threads to rotate and decreases the angle \u03b3, and the moment due to the force T\u03b8 works as a resistance. Here, the moment Mrot due to the force T around the intersection point A in the direction of angle \u03b3 for a single thread is determined as follows: Mrot \u00bc T\u03b8l0 cos \u03b3 Txl0 sin \u03b3: \u00f09\u00de If any energy losses due to these forces T\u03b8 and Tx are not considered and the effect of rubber's thickness is neglected, namely, by assuming Mrot \u00bc 0 in the actuator deformation and the radius R\u00bcRk, Eq",
            " Also, as for the item (3), a theoretical evaluation for the bulging deformation of the rubber tube should be needed. Moreover, as for the item (4), a theoretical modeling for the relative sliding distance between threads and the tube should be needed. In the following section, three kinds of deformation mechanisms are introduced, and the evaluation equation for the relationship between the axial load F and the pressure difference P\u2032 is proposed. Fig. 7(a) shows a schematic of the vertical cross section I-I (see Fig. 6). As shown in Fig. 7(a), the force T acts on the thread during rotation, and the S-shaped bending deformation can be obtained between the intersections of threads. The portion where threads contact mutually is shown by the shaded portion in the figure. The width of this contacting portion is denoted by ac and is given by the following equation: ac \u00bc a= sin 2\u03b3: \u00f010\u00de It can be found from Eq. (10) that although the distance l0 between intersections remains unchanged during deformation, the width ac of the contacting portion increases as the angle \u03b3 deviates from 451. Consequently, when the angle \u03b3 deviates from 451, higher strain energy is needed to maintain the bending deformation. Furthermore, the shear force Q shown in Fig. 6(a), which acts between the adjacent threads, increases due to the tensile force T, and a larger frictional force must be overcome to maintain the rotation of the threads. Therefore, in order to take the effect of the rotation of threads into consideration, both the rotational resistances due to the frictional force among threads, and the change in the bending strain energy associated with this rotational behavior should be considered. Based on the symmetry of the thread deformation shown in Fig",
            " These parameters can be determined from the following boundary conditions: yl \u00bc y\u2032l \u00bc 0 \u00f018\u00de at x\u00bc0, yl \u00bc yr \u00bc \u03b4=2 and y\u2032l \u00bc y\u2032r \u00f019\u00de at x\u00bcxE, yr \u00bc \u00f0\u03b4\u00feh\u00de=2 \u00f020\u00de at x\u00bcxO. Also, due to the equilibrium of moment for the part DA in a thread, the following equation can be derived. Q \u00f0xF xE\u00de \u00bc 2MA\u00feT\u00f0h\u00fe\u03b4\u00de: \u00f021\u00de The bulging behavior of the rubber through the rhombusshaped gap can be evaluated theoretically as a bulging problem, in which the rubber film is subjected to tensile stresses sx;rub and s\u03b8;rub on the film plane as shown in Fig. 6. The bulging deformation is caused by the pressure difference P\u2032 perpendicular to the film plane. Based on Hooke's law, the stresses sx;rub and s\u03b8;rub can be written by the following: sx;rub \u00bc Erub 1 \u03bd2rub \u00f0\u025bx\u00fe\u03bdrub\u025b\u03b8\u00de; s\u03b8;rub \u00bc Erub 1 \u03bd2rub \u00f0\u025b\u03b8\u00fe\u03bdrub\u025bx\u00de; 8>>< >>: \u00f022\u00de where \u025bx and \u025b\u03b8 denote strains along x- and \u03b8 directions, respectively, and given by the following: \u025bx \u00bc log L L0 \u00bc log cos \u03b3 cos \u03b30 ; \u025b\u03b8 \u00bc log D D0 \u00bc log sin \u03b3 sin \u03b30 : 8>>< >>: \u00f023\u00de The parameters Erub and \u03bdrub in Eq. (22) denote Young's modulus and Poisson's ratio of rubber, respectively. It is assumed that the bulging displacement u of the rubber film under the pressure difference P\u2032 takes the following form: u\u00bc C cos \u00f0xly ylx\u00de\u03c0 2lxly cos \u00f0xly\u00feylx\u00de\u03c0 2lxly ; \u00f024\u00de where parameters lx and ly in Eq. (24) can be given as follows: lx \u00bc l0 a sin 2\u03b3 cos \u03b3; ly \u00bc l0 a sin 2\u03b3 sin \u03b3; \u00f025\u00de and shown in Fig. 6. Also, the coefficient C in Eq. (24) can be derived from the following equation: \u222c\u03a9uP\u2032 dx dy\u00bc\u222c\u03a9 Nx \u2202u \u2202x 2 \u00feNy \u2202u \u2202y 2 \" # dx dy; \u00f026\u00de where the region \u03a9\u00bc f\u00f0x; y\u00dejjxj=lx\u00fejyj=lyr1g, and the forces Nx and Ny per unit length can be written in terms of sx;rub and s\u03b8;rub as follows: Nx \u00bc sx;rubtR; N\u03b8 \u00bc s\u03b8;rubtR: \u00f027\u00de Substituting Eq. (24) into Eq. (26), the parameter C can be written by the following: C \u00bc 32P\u2032 \u03c04 Nx l2x \u00feNy l2y \" #: \u00f028\u00de Firstly, the magnitude of relative sliding between the deformed tube and the surrounding threads is discussed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001107_2017-28-1939-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001107_2017-28-1939-Figure2-1.png",
        "caption": "Figure 2. Typical two-stage spur gearbox",
        "texts": [
            " The design variables vector X for two-stage gearbox is selected as, The range and type of the design variables used to solve the problem are shown in Table 2. Total Volume of the Gearbox The volume objective function is formulated considering the volume of gears, shafts, and gearbox frame. Thus, the volume function is defined as, (1) Volume of gears, (2) Volume of shafts, (3) Since the frame does not have a standard shape, the volume of frame Vfr is neglected. Hence the total volume constitutes the volume of the gears and the shafts. Figure 2 shows the gearbox arrangement considered for the optimization. In the equation 3, the length input and output shafts are taken as, (4) The second objective function of maximizing power includes the power losses in the gears, bearings, and seals. The split up of power losses in the gearbox according to [1] is as indicated in Figure 3. Therefore, in general form, the second objective function is written as, (5) Power losses in the gears PLgear are divided as load dependent and load independent losses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure6-1.png",
        "caption": "Fig. 6. DEM modeling with (a) rigid housing (b) flexible housing.",
        "texts": [
            " The apparent contact between a rolling element and a ring consists of n c micro-contacts. Each micro-contact i transmits a fraction qi of the undetermined radial force Q \u03c8 . In the radial direction, the radial load Q \u03c8 satisfies the following approximation: Q \u03c8 \u2248 i = n c \u2211 i =1 qi . n (14) where n is a unitary vector describing the main radial direction that connects the center of the inner ring with the center of the rolling element. The influence of deformable rings on the distribution of the mechanical load and stresses is investigated. Two systems are proposed in Fig. 6 . The inner ring is fixed on the shaft without clerance, represented by a discrete element with radius R inner . This mounting makes more stiff the inner ring and restricts its deformations. The system with rigid housing Fig. 6 (a) should provide a mechanical behavior close to the analytical approach with rigid rings usually assumed [11] but stresses within rings is now available. The assumption with half-clamping, Fig. 6 (b) allows to consider a flexible housing. Simulations are conducted with a time step t = 10 \u22128 s, due to the cohesive model of the rings. Considering the stiffness matrix given in expression (8) , where E \u03bcA \u03bc L \u03bc represents a typical stiffness, the time step with a safety coefficient of 1/10 is determined as follows : t = 1 / 10 \u00d7 ( \u221a m i E \u03bcA \u03bc L \u03bc ) . The radial load distribution Q \u03c8 and Von Mises stresses in rings are computed for differents radial loads F r = \u2212F r j with rigid and flexible housing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002772_j.ifacol.2016.07.125-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002772_j.ifacol.2016.07.125-Figure3-1.png",
        "caption": "Fig. 3. Positive angular displacement indicates that the fault is in the actuator 2, while negative indicates a fault in actuator 1.",
        "texts": [
            " / IFAC-PapersOnLine 49-5 (2016) 272\u2013277 The observed states are used to generate a residual signal given by: \u2016 r \u2016=\u2016 yr \u2212 y\u0302r \u2016 . (31) In fault free case the residual (\u2016 r \u2016) value is close to zero, while in faulty case the residual changes its value, bigger than a predefined threshold, as an indication of fault occurrence. When this change happens, the IMU measurements are considered to isolate the fault by taking in account the lectures of the roll displacement. To explain this idea, considers a VTOL in hover position as displayed in Fig 3, in this position the roll angle is equal to zero. If a fault occurs in actuator 1, the residual changes its value to one and the VTOL has a displacement in the counterclockwise direction. This effect is measured by the IMU as a negative roll displacement with respect to the x axis. Then by considering a predefined edge of the roll displacement, it is easy to identify the fault in the actuator one, as shown in Fig 3(a). On the other hand, if the roll displacement is positive, the fault is in the actuator two as shown in Fig. 3(b). Note that this idea is easy to implement in practice by measuring in real time the data given by the IMU. Moreover, if there is a rotational displacement in fault free case, the residual will be equal to zero, and therefore FDI is not performed, which clearly avoids false alarms. Finally, by using roll velocity, measured by the IMU, it is possible to isolate the actuator fault more rapidly than with roll displacement, even if the VTOL is not in hover position. 6. SIMULATION RESULTS The effectiveness of the cascade tracking controller, and the fault detection and isolation method using the qLPV representation applied to the nonlinear model (8) of the VTOL is presented on this section",
            " / IFAC-PapersOnLine 49-5 (2016) 272\u2013277 277 The observed states are used to generate a residual signal given by: \u2016 r \u2016=\u2016 yr \u2212 y\u0302r \u2016 . (31) In fault free case the residual (\u2016 r \u2016) value is close to zero, while in faulty case the residual changes its value, bigger than a predefined threshold, as an indication of fault occurrence. When this change happens, the IMU measurements are considered to isolate the fault by taking in account the lectures of the roll displacement. To explain this idea, considers a VTOL in hover position as displayed in Fig 3, in this position the roll angle is equal to zero. If a fault occurs in actuator 1, the residual changes its value to one and the VTOL has a displacement in the counterclockwise direction. This effect is measured by the IMU as a negative roll displacement with respect to the x axis. Then by considering a predefined edge of the roll displacement, it is easy to identify the fault in the actuator one, as shown in Fig 3(a). On the other hand, if the roll displacement is positive, the fault is in the actuator two as shown in Fig. 3(b). Note that this idea is easy to implement in practice by measuring in real time the data given by the IMU. Moreover, if there is a rotational displacement in fault free case, the residual will be equal to zero, and therefore FDI is not performed, which clearly avoids false alarms. - zq xq +f1 f2 (a) (b) threshold fault fault f2 f1 zq xq0 0 Fig. 3. Positive angular displacement indicates that the fault is in the actuator 2, while negative indicates a fault in actuator 1. Finally, by using roll velocity, measured by the IMU, it is possible to isolate the actuator fault more rapidly than with roll displacement, even if the VTOL is not in hover position. 6. SIMULATION RESULTS The effectiveness of the cascade tracking controller, and the fault detection and isolation method using the qLPV representation applied to the nonlinear model (8) of the VTOL is presented on this section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000081_12.2084733-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000081_12.2084733-Figure7-1.png",
        "caption": "Figure 7, Experimental setup for displacement measurement of curved unimorph DEA and corrugated DEA.",
        "texts": [
            " To produce the corrugated pattern, the positioning of the passive layer was alternated for each curved segment, as shown in Figure 6. The passive layer was placed on the concave surface of each curved segment to cause bending towards the concave side of the actuator and contraction of the corrugated structure. The corrugated DEA consisted of four VHB dielectric layers, and had a final width of 10mm, curved segments with 5mm radius of curvature and 90 degree subtended angle, and 15mm long straight segments. Actuation testing of the curved unimorph and corrugated DEA was performed using the setup in Figure 7. One end of the DEA was clamped in place with the width-axis held perpendicular to the ground to eliminate the effect of gravity on actuation. Driving voltage was applied to the DEA using alligator clips attached to a high voltage power supply. Actuation was observed by a CCD camera, and angle change of the curved unimorph and linear displacement of the corrugated DEA were measured in software. Proc. of SPIE Vol. 9430 943020-6 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/23/2015 Terms of Use: http://spiedl"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003363_1.g002683-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003363_1.g002683-Figure3-1.png",
        "caption": "Fig. 3 Approximate kinematics with flap angle \u03b2, lag angle \u03b6, and imposed lag-pitch coupling.",
        "texts": [
            " A physical interpretation of this result is simply that infinitesimal rotations commute, or that velocity vectors add. The exponential coordinates for the composite rotation dictated by the design geometry of Fig. 2 are expressed by 2 4 0 \u22121 0 3 5d\u03be1 2 4 sin \u03b4 0 \u2212 cos \u03b4 3 5d\u03be2 (1) For analysis, we re-parameterize the motion in terms of the canonical flap angle \u03b2 about an axis in the\u2212y\u0302 direction and lag angle \u03b6 about the \u2212z\u0302 direction, with both axes fixed in the hub frame. This arrangement is shown in Fig. 3. We separately impose a geometric lag-pitch coupling coefficient \u0394\u03b8\u2215\u0394\u03b6 tan \u03b4 and the resulting exponential coordinates for the composite rotation are 2 4 0 \u22121 0 3 5d\u03b2 2 4\u0394\u03b8\u2215\u0394\u03b6 0 \u22121 3 5d\u03b6 (2) The reparameterized expression encodes identical kinematics constraints as the original. The derived equation ofmotion, linearized for small deflections, will be equivalent if the flap and lag axes are coincident at eccentricity e and the rotational inertial of the small cross body and the blade about the pitch axis is neglected",
            " Instead of explicitly modeling two blades, the analysis is simplified by taking advantage of approximate symmetry and modeling only one blade and appropriately normalizing the hub inertia and motor torques by the number of blades. The limitations of this conventional technique when applied to our not precisely symmetric blades are discussed with the experimental results. The dynamics of the half propeller are developed as those of a three\u2013degree-of-freedom open-chain linkage with hub angle \u03c8 , lag angle \u03b6, and flap angle \u03b2 as defined in Fig. 3. The generic equation of motion is given by Eq. (3), where q f\u03c8 ; \u03b6; \u03b2g, a general result for open-chain dynamical systems [13]. The inertial matrix M q is a nonlinear function of the generalized coordinates and the Coriolis matrixC q; _q is a function of the coordinates and speeds. Both terms are derived directly from the kinematics depicted in Fig. 3 and inertial properties of the hub and blade body using the product of exponentials formula [13]. By convention, external (aerodynamic) forces applied to the rotor enter throughN and joint torques from the motor and hinge losses enter on the right as \u03c4motor and \u03c4hinge. M q q C q; _q _q N 1 Nb \u03c4motor \u03c4hinge (3) In anticipation of deriving a linearized governing equation, we identify a steady trim condition at rotor speed \u03a9 with coordinates q0 and velocities _q0 given by Eq. (4). Steady drag forces will cause the blades to lag backward by a small positive angle \u03b60 and lift forces will cause the blades to flap upward by a small positive coning angle \u03b20"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002660_978-3-319-06590-8_48-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002660_978-3-319-06590-8_48-Figure2-1.png",
        "caption": "Fig. 2 Rolling element bearing geometry [3]",
        "texts": [
            " This involved first identifying each impulse response with a user-set threshold, windowing the signal around this point, separating the step and impulse responses and then applying different signal processing techniques to each, with the intention of enhancing and properly locating each event. The semi-automated method applied pre-whitening and minimum entropy deconvolution to the signal before separation, and the respective events were then enhanced separately using complex Morlet wavelets. The time to impact was then calculated as the difference between the times corresponding to the peak of each enhanced event, and from this the spall size was estimated by assuming the exit impact occurs when the rolling element is halfway across the spall. Shown in Fig. 2 is the geometry of a rolling element bearing with an idealised rectangular outer-race fault. Sawalhi and Randall [3] showed that for \u03b1 = 0 the length of the spall\u2014subject to the above assumption\u2014is given by: lO \u00bc Tipfr D2 p d2 Dpfs \u00f01\u00de in which fr and fs are the shaft speed and sample rate (Hz), and Ti is the time to impact (samples). One potential limitation with the above approach is that while the fault exit event is usually characterised by a well-defined high-energy impulse response, the entry event is not always so clear",
            " A Br\u00fcel and Kj\u00e6r PULSE frequency analyser was used to record the data. Each test run was of 10 s duration, and a sample rate of 131,072 Hz was used. The tests were conducted over a range of speeds from 6 to 30 Hz in 6 Hz increments, with an applied radial load of 7 kN, which is 50 % of the rated dynamic load capacity of the test bearing. The test bearing used in this study was a Nachi 6205-2NSE9 single-row deepgroove ball bearing, containing nine balls of d = 7.94 mm diameter, with a bearing pitch diameter of Dp = 39 mm (see Fig. 2). To simulate a localised bearing fault, a notch of length lO = 1.6 mm was seeded on the bearing outer race using electric spark erosion. The notch was seeded across the entire race such that it was approximately 0.5 mm deep at the centre, sufficient to prevent the balls contacting the bottom of the notch. The fault was carefully positioned in the centre of the load zone for all the tests. The bearing with seeded fault is shown in Fig. 5. In order to develop an alternative spall size estimation methodology, all the fault signals over the speed range tested were observed closely, with a particular view to isolating the entry event"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001707_042148-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001707_042148-Figure1-1.png",
        "caption": "Figure 1. STC industrial robot and flexsplie",
        "texts": [
            " Published under licence by IOP Publishing Ltd There is essential difference between harmonic gear drive and traditional gear transmission. Harmonic gear drive mainly depends on the interaction between the flexspline and the wave generator, and the flexspline is one of the most critical components in the whole transmission system [1-3]. However, due to the constraints of operating conditions and materials processing, there is still an obvious disparity in the quality of flexspline. The principal part of the SHF-32-80 harmonic gear drive in STC industrial robot as is shown in Fig. 1. IMMAEE 2018 IOP Conf. Series: Materials Science and Engineering452 (2018) 042148 IOP Publishing doi:10.1088/1757-899X/452/4/042148 In practical application, the flexspline is subjected to transformation stress and cyclic loading, and is prone to fatigue failure. After use for a period of time, the flexspline can be damaged or even broken except for the oxidation and sticking of the surface of the gear teeth. These phenomena are not only detrimental to the smooth transmission of the system, but also cause production accidents"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003356_j.apm.2018.01.018-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003356_j.apm.2018.01.018-Figure2-1.png",
        "caption": "Fig. 2. 2D axisymmetric tyre model, including the location of the structural components and the regions modelled for the inner liner, sidewall, tread and rim strip rubber compounds.",
        "texts": [
            " Tyres are composite structures that consist of various rubber compounds and structural reinforcement. The rubber regions and structural components identified in the 23.5R25 earthmover tyre are shown in Fig. 1 . The rubber compounds include the inner liner, inner filler, soft apex, sidewall, shoulder, base and tread. The structural components include a radial steel body ply, five steel belt plies, and a bead bundle which consists of seventy-two individual wires. A 2D axisymmetric representation of the numerical model is shown in Fig. 2 . This model does not include all nine regions identified from the actual tyre cut-away, since some of the regions have similar mechanical properties and were grouped accordingly. In the 2D numerical model, the rubber compounds and bead wire were modelled using first order quadrilateral elements with a Full & Herrmann Formulation. These elements are formulated with displacement and pressure as variables to accommodate the incompressibility of rubber and to be compatible with the elements specified for the nonelastic components, such as the bead wire [13] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003410_j.matpr.2017.11.530-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003410_j.matpr.2017.11.530-Figure1-1.png",
        "caption": "Fig. 1: Types of cutting: (a) orthogonal cutting (b) Oblique cutting",
        "texts": [
            " Because of these difficulties alternative approaches developed as mathematical simulations where numerical methods are used. Among these numerical methods, finite element method is proved to be useful and widely used. Two types of cutting are used in analysis of metal cutting mechanics:  Orthogonal cutting  Oblique cutting. In orthogonal cutting, unwanted material is removed from the work piece by a cutting edge that is perpendicular to the direction of relative motion between tool and the work piece as shown in Fig. 1(a). In orthogonal cutting, the material removal process is assumed to be uniform along the cutting edge; therefore it is a two dimensional plane strain problem. In oblique cutting, the major cutting edge is inclined to direction of the cutting velocity with an inclination angle as shown in Fig. 1(b). Although most of the metal cutting operations are oblique, orthogonal cutting has been extensively studied because of its simplicity and giving good approximations [1]. Finite element method is basically defined as dividing a continuum system to small elements, describes element properties as matrices and assembles them to reach a system of equations whose solutions give the behavior of the total system. Different applications of FEM models for machining can be divided into six groups: a) tool edge design, b) tool wear, c) tool coating, d) chip flow, e) burr formation and f) residual stress and surface integrity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003363_1.g002683-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003363_1.g002683-Figure4-1.png",
        "caption": "Fig. 4 Power electronics, motor, and articulated hub for a 318-mmdiameter cyclic rotor.",
        "texts": [
            " Since the effective motor damping and stiffness coefficients cm and km are determined by choice of software speed control gains it is straightforward to conduct experiments at varied scales and speeds with identical governing equations. A 32-cm-diameter propeller embodying the kinematics of Sec. II was constructed. In combination with a commercial motor and custom electronic motor drive, the device allows controlled experiments of the cyclic system during which torque, speed, blade lag angle, and blade flap angle can bemeasured. The propeller shown in Fig. 4 is constructed from 3D printed plastic parts joined by simple stainless steel pins. The visible plastic screws serve only to retain the pins in place. A black commercial blade with an 11% thick, symmetric airfoil is bonded into the white custom blade grip. Two polytetrafluoroethylene (PTFE) plasticwashers in the lag hinge serve as thrust bearings to reduce the hinge friction under centrifugal loading. Critical rotor parameters for model calculations are summarized in Table 1. The eccentricity e of the hinge location was chosen to be 0",
            " The characteristic friction coefficients of the plastic\u2013plastic sliding contact and the silicone-lubricated steel\u2013plastic interfaces were estimated in separate tilted-plane tests. The motor and brushless motor controller drive the propeller rotation and are responsible for applying the once-per-revolution modulation of torque to excite the cyclic mode. The motor is a common brushless motor with a rotating shaft exposed at both ends. The motor orientation is directly measured by a contactless 4096- count magnetic rotary encoder mounted beneath the motor on the controller circuit board, shown in Fig. 4. This sensor observes the rotation of a diametrically polarized magnet bonded to the shaft end. These angle measurements are used to update the motor winding commutation at 40 kHz and update the speed controller at 2 kHz. In addition, this direct measurement of the hub rotation is used to calculate the modulation voltage ~V in order to ensure the phase and frequency of modulation remain synchronous with the hub rotation. The critical parameters for the electronic drive system are summarized in Table 2. The motor inertia, emf constant, and combined effective resistance of the motor and driver circuitry are fit values based on separate frequency response testing of the baremotor without an attached propeller. During experiments the rotor is supported on a vertical pylon to hold it out of ground effect. Thrust forces and reaction torques are measured by a small six-axis load cell atop the supporting pylon. The drive module of Fig. 4 containing the power electronics and motor is mounted directly to the load cell. The propeller is mounted to the rotating face of the motor by two screws to ensure that the propeller does not slip relative to the motor during testing. D ow nl oa de d by T U FT S U N IV E R SI T Y o n Fe br ua ry 1 5, 2 01 8 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .G 00 26 83 Experimentswere conducted atmean rotor speeds of 100, 200, and 300 rad\u2215s to determine the sensitivity of the cyclic response to drive amplitude inputs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003599_j.procir.2018.04.033-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003599_j.procir.2018.04.033-Figure11-1.png",
        "caption": "Fig. 11 Overhang structure",
        "texts": [
            " As a result, the hypothesis kept is a combination of the previous ones: the different contact area shape of the electron beam on the powder bed from the center to the border provides less energy to the powder. In addition, energy leaks near the border make the sintering of the powder weaker. Then, the rake passage might easily move the poorly sintered powder and the molten material. According to these investigations, the two main reasons of these defects are the loss of temperature and the contact of the rake with the specimen during the production. One of the most relevant advantages of the AM is the design freedom in geometry. However, the overhang part of a product, see Fig. 11, is one of the challengers to manufacture. Some geometric deformations occurs on the overhang zone during the build, as warping (or \u201ccurling\u201d depending on the authors), loss of thickness and loss of edge. Warping defects or curling defects corresponds to the curvature of the upper horizontal surface of an overhanging part, see Fig. 12 [25]. This effect has been observed in the existing bibliography and authors\u2019 experience. Warping defect is due to the thermal stress formed by the rapid solidification during the EBM process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002183_s00114-015-1259-6-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002183_s00114-015-1259-6-Figure8-1.png",
        "caption": "Fig. 8 The relationships between the shear force impulse, orientation angle of the feet, and the incline angle. a Changes in the shear force impulses and the orientation angle of gecko on different inclines. The solid lines with an arrow represent the shear force impulse acting on the front and hind feet. The dotted lines with arrows represent the orientation of the feet. \u03b2 is the propelling angle; \u03b3 is the orientation angle. b Overprinting the shear force impulses and the orientation angle of the gecko on different inclines. The variables on different inclines were distinguished by different colors",
        "texts": [
            " Changes in modes of propulsion with increased angle of the incline The change in the degree of inclines presents a great challenge for climbing animals. The climbing resistance increases with the increase of the component of gravity parallel to the surface of the incline (Dutto et al. 2004; Lammers et al. 2006). It is a major challenge for geckos to overcome this resistance and propel their body upwards. Measuring SRFs of geckos moving on different inclines can reveal the mechanisms by which the front and hind limbs overcome resistance and effect propulsion in response to the change of inclines (Fig. 8a). The fore-aft SRFs acting on the front and hind limbs show opposite trends while on inverted inclines (Fig. 5b), which indicates that the propulsion mechanism changes with the inclines. Additionally, the front limbs contribute more propulsion with increasing inclines. On horizontal substrates, gecko\u2019s front limbs generate a braking impulse and the hind limbs generate a propulsive impulse in the whole stance phase (Fig. 8a). This front limb deceleration and hind limb propulsion effectively cause resistance to external interference (Chen et al. 2006). The contribution of the front and hind limbs to propulsion on inclines differs among species and depends on the specific characteristics of the contact between the limbs and the surface (Dutto et al. 2004; Lammers et al. 2006; Lee 2011). On shallow inclines, the contribution of gecko\u2019s front limbs to propulsion increases gradually with the increase of angle\u2019s incline",
            " In contrast to the gecko, the frog cannot adhere to inverted inclines, which could be due to its wet adhesion system (Federle et al. 2006). Alternatively, this could also be due to the morphological structure of their hind limbs, which are incapable of the greater range of motion that would provide the necessary shear force opposite to the shear force of the diagonal front limb. As the angle of incline changes from 0 to 180\u00b0, the shear impulse acting on the hind limbs undergoes a turn backwards, as indicated by the solid lines in Fig. 8. The change in hind foot\u2019s orientation angle is not obvious, yet is indicated by the dotted line in Fig. 8. These results may be explained by the fact that the angles between each adjacent flexible toe are noticeably different from one another, as shown in Figs. 1c and 10b (Wang et al. 2010a). In other words, even without foot rotation, the direction of the SRF acting on the foot can be controlled by the angles between adjacent toes. The flexible fifth toe of the hind foot somewhat resembles a human thumb. This toe can construct a Y-configuration with the other toes (Fig. 10b), and the hind limb can further construct a Y- configuration by interlocking with the front limb",
            " Although SRFs acting on limbs can explain the majority of the functional differences between limbs, the influence of the morphology of the limbs, the contact model between the foot, and the surface on the SRFs are unclear. It is worth noting that the orientation angles of the feet change greatly with the propelling angles of corresponding feet when the geckos move on shallow inclines. Alternatively, the difference between the orientation and propelling angles decreases continuously as the incline becomes steeper (Fig. 8). These results may be related to the contact model between the foot and the surface when only the toes contact are in contact with it (adhesion contact) (Autumn et al. 2006a), the palm and toe are both in contact (dry friction and adhesion mixed contacts), or only the palm is in contact (dry friction contact). The supporting positive normal forces when climbing on shallow inclines are primarily generated by the palm of the foot, because the soft toes are of little value in generating a positive normal force impulse"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure6-1.png",
        "caption": "Figure 6: a) von-Mises stress and b) stress in xx-direction in sectional view 2 for the periodic lattice structure",
        "texts": [
            " The bearing and the force application are realized as described before. Stiffness. Figure 5 shows the simulation result for the z-displacement of the structure in figure 4. The displacement in the point of force application, which is the reference point to compare the different structures, is 0.142 mm. Together with the applied force of 300 N, a stiffness of 2117 N/mm results for the structure. To make the performance more comparable to other structures, the ratio of stiffness to mass is calculated. Here, a value of 14.4 N/(mm*g) arises. Maximum Stress. In figure 6 a), the von-Mises comparison stress is depicted for the regular lattice structure under the given load. The maximum value for the von-Mises stress appearing in the simulation is 50.58 N/mm\u00b2. Under the assumption, that this maximum stress linearly depends on the applied force, a maximum load of 406.6 N can be calculated for the obtaining of the limit of elasticity. Hence, the ratio of this force to the part\u2019s mass is 2.76 N/g. This value will be used in order to compare the structure with the further ones in the following sections. In figure 6 b), the sectional view at cutting plane 2 in figure 3 is depicted. Thereby, the stress in xx-direction for the cross section of six central struts in this plane is investigated. It can be recognized, that in the middle of the struts, only minor stresses appear. In contrast, in the upper sections pull forces and in the lower sections push forces can be observed, whereat both have similar absolute values. This leads to the conclusion, that there exists a major bending load on the horizontal struts",
            " Due to the extensive reduction of the structure\u2019s mass, the ratio of stiffness to mass improves for 81 % towards a value of 25.9 N/(mm*g). Maximum Stress. In figure 10, the von-Mises comparison stress and the stress in xx-direction in cutting plane 2 are depicted for the flux of force adapted lattice structure. Figure 10 b) shows the stress in xx-direction for the cross section of the struts in cutting plane 2 (see figure 3). In order to make the results more comparable, the same colour scale has been used as in figure 6 b). It can be recognized, that only minor stress variations appear inside the single struts. This leads to the conclusion, that the bending loads in this sectional view have severely been reduced. Nevertheless, the maximum value for the von-Mises stress has risen to 74.63 N/mm\u00b2 (see figure 10 a). Under the assumption, that this maximum stress linearly depends on the applied force, a maximum load of 221.1 N can be calculated for the obtaining of the limit of elasticity. This means a worsening of 46 % compared to the periodic structure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000652_lmag.2012.2214027-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000652_lmag.2012.2214027-Figure1-1.png",
        "caption": "Fig. 1. Magnetic potential contour plot with superimposed field lines (dashed) of a dipole with |m| = 18 A\u00b7mm2, oriented at 45\u25e6, located 1 mm below a soft magnetic half-space of permeability (a) \u03bcr = 5 and (b) \u03bcr = 100 (b). The potential contour labels are given in milliampere.",
        "texts": [
            ", Jackson [1999]): a point charge (or dipole) in proximity of a dielectric half-space of permittivity \u03b5 generates an image charge (or dipole) located symmetrically; the field within the dielectric is that generated by a second image charge (or dipole) at the same position of the source. The magnitudes of the image charges (or dipoles) are given exactly by the scaling factors that appear in (7) with the substitution \u03bcr\u2192\u03b5r . If we plot the solution for a magnetic dipole in proximity of a permeable half-space with \u03bcr = 5 and \u03bcr = 100, we find the familiar trend of the potential contour lines that are progressively forced to close in vacuum and avoid penetrating the soft medium as shown in Fig. 1. When the permeability is very large, a soft material behaves with magnets as a metal does with charges: the (H or E) field lines are forced to impinge perpendicularly to the boundary, and the potential inside the material vanishes, with only surface charges remaining that allow for the boundary conditions to be satisfied. Having demonstrated the validity of the image method for the problem at hand, and considering that a uniformly magnetized body is a collection of aligned magnetic dipoles, the effect of the permanent magnet on a soft half-space with high permeability (\u03bcr 1) is to create a mirror-symmetric version of the magnetized body"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001104_s11665-017-2607-9-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001104_s11665-017-2607-9-Figure2-1.png",
        "caption": "Fig. 2 3D configuration and geometry of manufactured specimens with biomimetic surface: (a) convex, (b) deep concave, (c) shallow concave",
        "texts": [
            "06 mm was employed to prepare the biomimetic surfaces on specimens. The treated surface of the specimens was engineered to imitate the surface morphology of earthworm covered by some concave and convex spots. This feature is believed to provide excellent adaptability in their corresponding habitats (Ref 18). Since the biomimetic laser technique was described elsewhere (Ref 19, 20), it is only briefly introduced here. The laser parameters used for manufacturing are given in Table 1, and the surface morphologies of manufactured specimens are shown in Fig. 2. The details of 3D profile of convex, deep concave and shallow concave on the specimens are depicted in Fig. 2(a), (b) and (c), respectively. Microhardness along the depth of transverse section of units was measured using Vickers hardness tester with 200 g load and 15 s loading time (HV-1000,Shanghai Precision Instruments Cooperation). The microstructures of the biomimetic unit were analyzed by scanning electron microscopy (SEM, Zeiss EVO 18 special edition) technique. The friction and wear experiments were carried out on the MG-2000 pin-on-disk sliding wear test rig with applied loads of 100 N. The wear experiments were performed at the constant rotating speed of 300 rpm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001969_polb.23930-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001969_polb.23930-Figure2-1.png",
        "caption": "FIGURE 2 Lamella misaligned in the x\u2013z plane, so that / 5 2w 1 h.",
        "texts": [
            " Figure 1(b) shows the resulting lamellar reflection in reciprocal space and the equivalent symmetrical JOURNAL OF POLYMER SCIENCE WWW.POLYMERPHYSICS.ORG FULL PAPER WWW.MATERIALSVIEWS.COM JOURNAL OF POLYMER SCIENCE, PART B: POLYMER PHYSICS 2016, 54, 308\u2013318 309 reflections that together make up the four spot pattern. The radial distance from the origin to the reflection is (proportional to) 1/t while the x and z components are sx5 sin/ t 5 sin/ L1a\u00f0 \u00decosw 5 sin h2w\u00f0 \u00de L1a\u00f0 \u00decosw sz5 cos/ t 5 cos/ L1a\u00f0 \u00decosw 5 cos h2w\u00f0 \u00de L1a\u00f0 \u00decosw (1) Figure 2 shows the case of misorientation in the x\u2013z plane, where h 6\u00bc 0 and / 5 h \u2013 w. In Figures 1(a) and 2, L is the molecular length within the crystalline lamella, and a is the thickness of the amorphous layer alongm. The repeat distance or \u201clong period\u201d is t5 (L1 a) cosw. For the nylon fiber used as an example here a is L/2, while for the liquid crystalline materials a is small. In mechanical terms, a change in h is a body rotation of the lamella, while a change in w is a shear. Thus, a large range FIGURE 3 predicted diffraction spot shape depending on the relative range of h and w",
            " The relation between h and / is FIGURE 7 Angular relations when m does not lie in the x\u2013z plane. h, /, and b are measured from the origin of the sphere (radius r). n is measured around a circle on the surface of the sphere. FIGURE 6 A reflection broadened in x and y, shown in the z 5 0 plane. The reflection comes from a unit rotated by an angle g from the exact diffraction position. FULL PAPER WWW.POLYMERPHYSICS.ORG JOURNAL OF POLYMER SCIENCE 314 JOURNAL OF POLYMER SCIENCE, PART B: POLYMER PHYSICS 2016, 54, 308\u2013318 no longer as shown in Figure 2, because m need not be in the x\u2013z plane. Figure 7 shows h, /, and b, the angle between them. Following Fraser48 page 88, consider all units with m at an angle b to the vector (r, /); as they are rotationally averaged, each of these units contributes a known intensity I(r,b) to the diffraction at r, /. The axes m of all these units lie on a cone of semiangle b, which intersects the surface of the sphere of radius r at a circle, radius sin b as shown in Figure 7. At each value of n, the rotation angle around this circle of radius sin b, the number of diffracting units is given by G(h)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000169_ext.12016-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000169_ext.12016-Figure14-1.png",
        "caption": "Figure 14 Relation between angle \u03c8 and location z of a transverse section at helical gearings.",
        "texts": [
            " The same is considered for the tangential displacements at the flute for torque transmission. As a result of the forming process, simulation contact forces can be identified for every time increment, respectively, tool rotation angle. For adjacent time increments, the nodal contact forces and node locations are selected and stored. From the current processing time and rotational tool velocity, the tool rotation angle \u03c8 can be calculated. This calculated angle of the 2D model corresponds with one transverse section of the 3D model. Figure 14 together with Eq. 2 illustrate this relation. z (\u03c8) = d 2 \u00b7 sin \u03c8 (t) tan \u03b2 (2) Thus, the contact nodes are located in a 3D coordinate system. A new mesh, consisting of shell elements is generated with this information. The stored nodal contact forces are applied as boundary conditions to this new mesh, considering the normal direction of the tool gearing and distance between adjacent nodes in axial direction. The shell elements and their nodal forces represent the loading for the linear elastic model of the tool"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002577_fpmc2015-9540-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002577_fpmc2015-9540-Figure4-1.png",
        "caption": "Figure 4 \u2013 Free body diagram of the load acting on a gear: Fload is in this case representative of the total load acting on the gear and Fj is the reacting force of the JB while dj is the lever arm of Fj",
        "texts": [],
        "surrounding_texts": [
            "relevant for sudden variations of the load applied on the JB, such as those related to delivery pressure ripple in pumps. To overcome these limitations, this work presents the coupling of a CFD model with a rigid body motion model of the gears based on the balance of the forces applied to it. Because of its integration within HYGESim, this model represents the first effort of this kind in the evaluation of EGM documented in literature. The structure of the following part of the paper is as follows: in section 2 the details of the evaluation of the load acting on the gears shafts due to fluid pressure are provided. Section 3 gives the details of the CFD model and its solver formulated by the authors specifically for EGMs. Section 4 an initial validation procedure of the model is explained. In section 5 the algorithm for the evaluation of the micro-motions is shown. Section 5 provides significant results, along with a first validation based on experimental results on a commercial pump.\n2. EVALUATION OF LOADS APPLIED ON THE JOURNAL\nThe lumped parameter model of HYGESim is utilized to determine all the instantaneous forces acting on each gear. In the radial plane, these forces are shown in Fig. 3. This evaluation is made according to the procedure described in [22] and here summarized for the sake of clarity. The resultant radial force and torque acting on the gear 2 (free gear) is made separately considering the x and y forces. In particular, the x and y components of the pressure force of each TSV is evaluated by the fluid dynamic module of HYGESim, by\ndetermining also all the TSV projection areas xz and yz as shown in Fig. 3.\n\ud835\udc53\ud835\udc65\u20d7\u20d7  \u20d7(\ud835\udf17) = \ud835\udc5d\ud835\udc56(\ud835\udf17) \u22c5 \u03a9\ud835\udc66\ud835\udc67(\ud835\udf17)\ud835\udc56 \ud835\udc65 (1a) \ud835\udc53\ud835\udc66\u20d7\u20d7  \u20d7(\ud835\udf17) = \ud835\udc5d\ud835\udc56(\ud835\udf17) \u22c5 \u03a9\ud835\udc65\ud835\udc67(\ud835\udf17)\ud835\udc56 \ud835\udc66 (1b)\nThe forces acting the TSVs of each gear are then summed up to obtain the resultant instantaneous force:\n\ud835\udc39\ud835\udc5d,\ud835\udc65 \ud835\udc54\u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 (\ud835\udf17) =\u2211\ud835\udc53\ud835\udc65,\ud835\udc56 \ud835\udc54 (\ud835\udf17)\n\ud835\udc56\n(2a)\n\ud835\udc39\ud835\udc5d,\ud835\udc66 \ud835\udc54\u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 (\ud835\udf17) =\u2211\ud835\udc53\ud835\udc66,\ud835\udc56 \ud835\udc54 (\ud835\udf17)\n\ud835\udc56\n(2b)\nThe line of action of the resultant force \ud835\udc39\ud835\udc5d,\ud835\udc65 \ud835\udc54 is evaluated considering the distances from the x axis (Fig. 3) of each \ud835\udc53\ud835\udc5d,\ud835\udc65:\n\ud835\udc4c\ud835\udc54(\ud835\udf17) = \u2211 (\ud835\udc53\ud835\udc65,\ud835\udc56\n\ud835\udc54 (\ud835\udf17) \u22c5 \ud835\udc66\ud835\udc56 \ud835\udc54 (\ud835\udf17))\ud835\udc67 \ud835\udc56=1\n\ud835\udc39\ud835\udc5d,\ud835\udc65 \ud835\udc54 (\ud835\udf17)\n(3a)\nSimilarly for the \ud835\udc39\ud835\udc5d,\ud835\udc66 \ud835\udc54 :\n\ud835\udc4b\ud835\udc54(\ud835\udf17) = \u2211 (\ud835\udc53\ud835\udc66,\ud835\udc56\n\ud835\udc54 (\ud835\udf17) \u22c5 \ud835\udc65\ud835\udc56 \ud835\udc54 (\ud835\udf17))\ud835\udc67 \ud835\udc56=1\n\ud835\udc39\ud835\udc5d,\ud835\udc66 \ud835\udc54 (\ud835\udf17)\n(3b)\n\ud835\udc4c\ud835\udc54 (\ud835\udf17 ) and \ud835\udc4b\ud835\udc54(\ud835\udf17) are used as instantaneous arms respectively of the \ud835\udc39\ud835\udc5d,\ud835\udc65 and \ud835\udc39\ud835\udc5d,\ud835\udc66 to determine the instantaneous moments acting on the gear shaft due to fluid pressure. The torque due to fluid pressure on the free gear (gear 2) is assumed to be instantaneously balanced by the contact force Fc. This assumption permits to explicit the instantaneous Fc. Once Fc is known, the torque \ud835\udc47\ud835\udc61\ud835\udc5c\ud835\udc61 applied on external shaft connected to gear 1 (input torque, in case of pump, output torque, in case of motor) can be determined by the contributions of this contact force and the moments due to the pressure forces \ud835\udc39\ud835\udc5d,\ud835\udc65 \ud835\udc541 and \ud835\udc39\ud835\udc5d,\ud835\udc66 \ud835\udc541 . For each gear, the total radial load is given by:\n\ud835\udc39 \ud835\udc59\ud835\udc5c\ud835\udc4e\ud835\udc51 \ud835\udc54 = \ud835\udc39\ud835\udc5d\ud835\udc65 \ud835\udc54\u20d7\u20d7 \u20d7\u20d7  \u20d7 + \ud835\udc39\ud835\udc5d\ud835\udc66 \ud835\udc54\u20d7\u20d7\u20d7\u20d7\u20d7\u20d7 + \ud835\udc39\ud835\udc50 \ud835\udc54\u20d7\u20d7\u20d7\u20d7  \u20d7 (4)\nSince in EGMs each gear is supported by two JBs, it can be assumed that this load acting on each gear is equally shared between two JBs (Fig, 4). Making the realistic assumption of symmetric journal bearings and that no external radial forces are acting on the shafts the equilibrium of each shaft leads to\n{ \ud835\udc39\ud835\udc3f\ud835\udc5c\ud835\udc4e\ud835\udc51\u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 + \ud835\udc39\ud835\udc571\u20d7\u20d7\u20d7\u20d7  \u20d7 + \ud835\udc39\ud835\udc572\u20d7\u20d7\u20d7\u20d7  \u20d7 = 0\n\ud835\udc39\ud835\udc571\u20d7\u20d7\u20d7\u20d7  \u20d7 \u22c5 \ud835\udc51\ud835\udc571 + \ud835\udc39\ud835\udc572\u20d7\u20d7\u20d7\u20d7  \u20d7 \u22c5 \ud835\udc51\ud835\udc572 = 0 (5)\n\ud835\udc39\ud835\udc571\u20d7\u20d7\u20d7\u20d7  \u20d7 = \ud835\udc39\ud835\udc572\u20d7\u20d7\u20d7\u20d7  \u20d7 = \u2212 \ud835\udc39\ud835\udc3f\ud835\udc5c\ud835\udc4e\ud835\udc51\u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7\n2 (6)\nIn this work, shaft deformation is not considered and consequently the gear axis will not tilt during the operation. Therefore, for each gear the position of the corresponding shaft inside the bearing can be considered function of only two\n3 Copyright \u00a9 2015 by ASME\nDownloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/86536/ on 03/17/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
            "variables such as the two eccentricities shown in Fig. 5. Furthermore, in this analysis the bearings are considered to be rigidly connected to the sliding bushing blocks of Fig. 2. As described in [21] and [22], the sliding bushing blocks can have relative motion inside the pump casing according to the existing clearance between the two elements. In case of non-pressure compensated EGM, the bearings are assumed rigidly connected with the EGM case, thus in a fixed position in space.\n3. CFD MODEL OF THE JOURNAL BEARING Since each gear is rigidly connected to its shaft, in order to find the displacement of the gears with respect to the pump case (thus finding the features of the radial gap), it is necessary to calculate the eccentricity of the journal with respect to the bearing. This can be done solving the pressure field in the lubricating gap between the two JB surfaces. This can be accomplished by solving the Reynolds equation. Under the full film lubrication assumption, a modified version of the Reynolds equation has been derived ad-hoc for the current case:\n\u2207 \u22c5 (\u2212\n\ud835\udf0c\u210e3 12\ud835\udf07 \u2207\ud835\udc5d)\n\u23df 1\n+ \ud835\udf0c\u2207\u210e\n2 (\ud835\udc63\ud835\udc61)\u23df 2 \u2212 \ud835\udf0c\ud835\udc63\ud835\udc61 \u22c5 \u2207\u210e\ud835\udc61\u23df 3\n+ \ud835\udf0c \ud835\udf15\u210e\n\ud835\udf15\ud835\udc61\u23df 4\n= 0 (7)\nwhere \u210e = \u210e\ud835\udc61 \u2212 \u210e\ud835\udc4f.\nEq. (7) assumes a null velocity for the surface of the bearing to and neglects the spatial derivative terms of the surface velocities. Four main terms can be identified from Eq. (7): the Poiseuille term (1), that represents the diffusion of pressure from the boundaries; the wedge term (2), that accounts for the pressure generations due to variations in the gap height; the translation squeeze term (3), whose action results from the translation of an inclined surface and finally the squeeze term (4), which predicts the pressure generation due to the reciprocal movement of the two surfaces towards or away from each other. The terms (2), (3) and (4) are associated to the hydrodynamic pressure which is the pressure generated by the reciprocal movement of the components.\n4 Copyright \u00a9 2015 by ASME\nDownloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/86536/ on 03/17/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
            "Analytical solutions for Eq. (7) in steady-state conditions exist for the JB problem such as the Sommerfeld solution for infinitely wide journal bearings, or the Ocvirk and Du Bois for narrow bearings ( \ud835\udc59\n\ud835\udc51 < 0.5). However, since the solution strongly depends\non the geometrical boundaries of the problem, the application of these analytical solutions is unsuitable for cases such as EGMS for which the diameter to width ratio of the bearing approaches the value of 1. A solution for these cases can be obtained by a numerical approach. Since pressure variation along the gap height are neglected in the Reynolds equation, the lubricating interface can be discretized with a 2D mesh. In this work, the Eq. (7) is then solved though a finite volume method, using a preconditioned gradient algorithm with a diagonalized incomplete Cholesky preconditioner [25]. Thanks to the simple geometry of the domain (as shown in Fig. 6) a structured mesh, preferred in finite volume methods, has been used. This allows for a fairly small number of fluid cells. The curvature of the film can be neglected being the gap height much smaller than all the other dimensions of the domain. With that, the fluid mesh can be unwrapped to obtain a more convenient Cartesian reference system, instead of a cylindrical one. The gap height \u2013 the necessary input to solve the fluid pressure according to Eq. (7) \u2013 can be defined with reference to the reference system of Fig. 5 as:\n\u210e(\ud835\udf03) = \ud835\udc50 \u2212 \ud835\udc52\ud835\udc65 sin ( \ud835\udf03\n\ud835\udc5f ) \u2212 \ud835\udc52\ud835\udc66 cos (\n\ud835\udf03 \ud835\udc5f ) (8)\nWhile, as concerns the surface speed of the journal:\n\ud835\udc63\ud835\udc61 = ?\u20d7? \u22c5 \ud835\udc5f (9)\nwhere 0 \u2264 \ud835\udf03 < 2\ud835\udf0b \u22c5 \ud835\udc5f. Radial loads acting on the shaft can be supported by eccentric motion position of the journal which leads to hydrodynamic pressure generation in the region of minimum film thickness. This pressure can exceed 10 MPa in high pressure EGM JBs. For this reason, changes in the fluid properties due to fluid pressure need to be taken into account. In the proposed model,\nthis is due by considering a variable density and viscosity according to [26]:\n\ud835\udf0c = \ud835\udf0c0[1 + \ud835\udefd\ud835\udc5d(\ud835\udc5d \u2212 \ud835\udc5d0)]\n\ud835\udf07 = \ud835\udf070\ud835\udc52 \ud835\udefc\ud835\udc5d\nThese relations have been already implemented in [23,24,27]. Cavitation phenomena are modeled in a simplified fashion, by limiting the minimum pressure value to the oil saturation pressure. This assumption was also made in the study of other lubricating interfaces of positive displacement machines, as reported in [23,24,27].\n4. EVALUATION OF GEAR MICRO MOTIONS\nThe complete algorithm used to solve the Reynolds eq. along with the radial micro-motions of the gears is shown in the flow chart of Fig. 7. The algorithm was conceived for solving the case of dynamic external loads, so that the instantaneous JB load can be taken into account also as concerns its rate of change. At first, an initial value for the eccentricities for each JB is assumed as well as the two initial squeeze velocities. The Reynolds eq. is then solved and the pressure distribution evaluated. Once the pressure field is known, a numerical integration of it over the spatial domain is performed to calculate the forces arising from the gap, as:\n\ud835\udc39\ud835\udc65\ud835\udc43 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 = \u2211\ud835\udc5d\ud835\udc56 \u22c5 \ud835\udc46\ud835\udc56 \u22c5 cos ( \ud835\udc65\ud835\udc56 \ud835\udc5f )\n\ud835\udc5b\n\ud835\udc56=1\n(10a)\n\ud835\udc39\ud835\udc66\ud835\udc43 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 = \u2211\ud835\udc5d\ud835\udc56 \u22c5 \ud835\udc46\ud835\udc56 \u22c5 sin ( \ud835\udc65\ud835\udc56 \ud835\udc5f )\n\ud835\udc5b\n\ud835\udc56=1\n(10b)\nThe force balance of the gear of Fig. 3 has to be fulfilled in order to sustain the external load applied:\n{ \ud835\udc39\ud835\udc65\ud835\udc3f\ud835\udc5c\ud835\udc4e\ud835\udc51\u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 + \ud835\udc39\ud835\udc65\ud835\udc43\u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 = 0\n\ud835\udc39\ud835\udc66\ud835\udc3f\ud835\udc5c\ud835\udc4e\ud835\udc51\u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 + \ud835\udc39\ud835\udc66\ud835\udc43\u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 = 0 (11)\nAs shown in Fig. 7, if the condition of static equilibrium of Eq. (11) is not respected, a \u201cforce balance loop\u201d is involved to iterate around the hydrodynamic squeeze term of the Reynolds equation, thus affecting the pressure field in the JB gap. In particular, the squeeze velocity \ud835\udf15\u210e\n\ud835\udf15\ud835\udc61 , evaluated as function of the\ntwo eccentricities \ud835\udc51\ud835\udc52\ud835\udc65 \ud835\udc51\ud835\udc61 , \ud835\udc51\ud835\udc52\ud835\udc66 \ud835\udc51\ud835\udc61 , is calculated using the multidimensional root-finding Powell\u2019s Hybrid method. Once the force balance is fulfilled, the squeeze term can be integrated through a simple Euler integration, and the film thickness is then updated according to the new eccentricities values. At this point, the simulation can then proceed to the next time step.\n5 Copyright \u00a9 2015 by ASME\nDownloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/86536/ on 03/17/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use"
        ]
    },
    {
        "image_filename": "designv11_13_0001136_s00170-017-0625-2-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001136_s00170-017-0625-2-Figure3-1.png",
        "caption": "Fig. 3 Scanning strategy of the thin-wall structure. The deposition directions are indicated by arrows",
        "texts": [
            " The pulse repetition frequency is 3 Hz with 4 ms as the width of the pulse (Fig. 2a). The average energy of single-shot pulsed laser measured with a commercial joulemeter is 38 J. Based on our experience, using a negative defocus laser beam as an energy source is a better method for attaining a smooth side surface. The smoothening of the side surface is a result of the action of gravity and surface tension owing to the remelting on the side of the deposition under the negative defocused laser beam (Fig. 2b). Figure 3 shows the scanning strategy of the thin-wall structure, in which the deposition directions are indicated by arrows. Eight different thin-wall samples were deposited with 1, 2, 3, 5, 8, 13, 21, and 34 layers, which are referred to as S1, S2, S3, S5, S8, S13, S21, and S34, respectively. The processing parameters for the thin wall sample are recorded in Table 2. Moving distance of the Z-axis in each layer in S1\u2013S34 samples is recorded in Table 3. In this study, the shapes of molten pools vary with the deposition process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001853_s10846-015-0233-z-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001853_s10846-015-0233-z-Figure13-1.png",
        "caption": "Fig. 13 Coordinates definition for defining D-H parameters presented in Table 2",
        "texts": [
            " 10, O is the position vector from the origin of the general coordinate system, G, to the local coordinate system of the platform, B, si1 represents the position of ith hip joint in B, ui is the ground contact point of the ith leg in G, li1 is the ith hip leg vector in G, R \u2208 so(3) is the rotational matrix for the main body of the robot B with respect G, and si2 is calculated as follows: si2 = \u23a1 \u23a3 si1x + (\u22121)i \u00b7 l1i \u00b7 cos(\u03b1i) si1y + (\u22121)i \u00b7 l1i \u00b7 sin(\u03b1i) si1z \u23a4 \u23a6 (12) If changing the foot contact locations is preferred within the controller, the same equations for \u03b1i , \u03b2i , and \u03b3i as presented in Fig. 10 can be used where ui , the foot tip vector, is given. Each leg follows a planned trajectory to reach the new foothold [28]. The desired position and velocity of the changing foot is used to find the required joint positions and speeds of the leg based on the kinematics of the leg. In this analysis, there are three different coordinate systems: {b} which is the body fixed frame, {g} which is the ground (inertial) coordinate system, and {f } which is the foot coordinate system as shown in Fig. 13. According to the physical parameters of the robot, the desired trajectory, shown in Fig. 12, has been chosen for each leg while changing the location of the foothold. Considering this trajectory, the following can be written for the position and velocity of the legs: [ xf/b(t) zf/b(t) ] = [ xf/g(t) zf/g(t) ] \u2212 [ xb/g(t) zb/g(t) ] (13) and [ x\u0307f/b(t) z\u0307f/b(t) ] = [ x\u0307f/g(t) z\u0307f/g(t) ] \u2212 [ x\u0307b/g(t) z\u0307b/g(t) ] (14) and [ x\u0308f/b(t) z\u0308f/b(t) ] = [ x\u0308f/g(t) z\u0308f/g(t) ] \u2212 [ x\u0308b/g(t) z\u0308b/g(t) ] (15) The right hand side of Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003073_060201-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003073_060201-Figure3-1.png",
        "caption": "Fig. 3. (color online) Working principle of the induction melting.",
        "texts": [
            " An attempt is made to understand the complex process and describe the intrinsic physical properties in consecutive induction melting by numerical simulation as discussed in the next sections. 060201-2 When the slowly rotated and fed superalloy electrode was lowered into the conical induction coil with alternating current, a time-varying current could induce a varying magnetic field. In this setup, an equal but opposite electric current could be generated on the surface of the superalloy electrode, the surface of the electrode will be quickly heated up to its melting point. Figure 3 shows the schematic of the induction melting process. The melting physical principle was based on the well-known Joule-Lenz law: Q = I2Rt, (1) where I is the effective current (in unit A), R the resistance of the metal piece (superalloy electrode) (in unit \u2126), and t the melting time (in unit s). In the induction melting process, the control of the rotating and feeding electrode was an important issue. Factors including electrode vertical feed rate, melting temperature and thermodynamic parameters are described by the continuity, momentum and energy equations as follows: \u2207(\u03c1u) = 0, (2) \u2202 (\u03c1u) \u2202 t +\u2207(\u03c1uu) =\u2212\u2207p+\u2207(\u00b5\u2207u)+\u03c1g[1\u2212\u03b2 (T \u2212Tref)], \u03c1Cpu \u00b7\u2207T = k\u2207 2T +Q, (3) \u2207 = \u2202 \u2202x \ud835\udc56+ \u2202 \u2202y \ud835\udc57+ \u2202 \u2202 z \ud835\udc58, (4) u = ux\ud835\udc56+uy\ud835\udc57+uz\ud835\udc58, (5) where \u03c1 is the liquid metal flow density (in units kg/m3), u the electrode vertical feed rate (in units m/s), p the ambient pressure (in unit Pa), T the free surface temperature (in unit \u25e6C), Tref the reference temperature (in unit \u25e6C), Q the inductive heating energy (in unit J), k the thermal conductivity (in units W/(m\u00b7K)), \u03b2 the dynamic viscosity (in units m2/s), Cp the specific heat capacity (in units J/(kg\u00b7K)), and the Hamiltonian operator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000748_tie.2013.2276025-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000748_tie.2013.2276025-Figure10-1.png",
        "caption": "Fig. 10. Magnetic flux density B in the system for nnom = 1420 r/min.",
        "texts": [
            " 5) The principal parameters of the computation are as follows: a) magnetic field: 16060 second-order elements, 32023 DOFs; b) temperature field: 736 second-order elements, 1569 DOFs; c) number of threads: 4 (two physical and two logical); d) process memory: 315 MB; and e) elapsed time for one example: 00:31.784 s. Even for the above simplifying assumptions that reduced the time of computations more than ten times, the results exhibit a very good accordance with measurements, which is discussed in Section V-C. Both computations and measurements provided many results. In the next paragraphs, we will show and discuss the most important ones. Fig. 10 shows the distribution of magnetic flux density B in the system for the nominal revolutions of the billet nnom = 1420 r/min. Only very small areas near the corners of joints between the permanent magnets and the magnetic circuit are oversaturated. Analogously, Fig. 11 shows the distribution of the volumetric Joule losses pJ produced by the currents induced in the rotating billet at the same revolutions. In accordance with theory, its highest values are generated in the surface layers of the billet"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003666_access.2018.2879649-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003666_access.2018.2879649-Figure2-1.png",
        "caption": "FIGURE 2. Structure of Scale Cessna 182 [10], [11].",
        "texts": [
            " UAV SELECTION In this paper, the scale model of the Cessna 182 is selected as the experimental UAV, and we call this selected UAV \u2018\u2018scale Cessna 182\u2019\u2019 by making its clear definition in order to distinguish it from the prototype Cessna 182. The ratio of the scale Cessna 182 to the prototype Cessna 182 is 1: 6.65. Cessna series are manufactured in Wichita, Kansas, USA, and they are widely used by the US Air Force for education and training purposes. Because the performance of the scale machine is similar to the prototype, the scale model is often used to carry out relevant experiments using the real flight equipment. The scale Cessna 182 is shown in FIGURE 1 [9], and the structure of the body can be seen in FIGURE 2 [10]. The geometric parameters of the scale Cessna 182 are also shown in FIGURE 3 [11] and Table 1 [12]. At present, the research of UAVs is divided into two types. The first category is the application research of UAV. For example, after UAVs are loaded with different payloads to collect information, subsequent data processing is performed with UAVs equipped with camera devices or sensors, image VOLUME 6, 2018 70735 detection and environmental exploration work; the second type of research is UAV motion analysis, improvement or controller design"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003923_0954406219850855-Figure14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003923_0954406219850855-Figure14-1.png",
        "caption": "Figure 14. The machined herringbone grooves on the shaft (No. 1).",
        "texts": [
            " In terms of the herringbone groove geometrical parameters of the sensitivity analysis, 11 groups of herringbone groove parameters are designed to experimentally investigate the influence of herringbone grooves on the radial runout of the ultra-high-speed aerostatic spindle, as shown in Table 3. If the results of sensitivity analyses, which are obtained based on the proposed improved FEM model, are in accordance with those acquired by experimental tests under the same operating conditions, the improved FEM model will be indirectly validated. The herringbone grooves of two aerostatic journal bearings are directly machined on the shaft of the spindle, as shown in Figure 14. A radial runout test bench of the herringbone grooved ultra-high-speed aerostatic spindle is depicted in Figure 15(a). The speed range of the ultra-high-speed aerostatic spindle used in the experiments is from 20,000 r/min to 200,000 r/min. The spindle radial runout is measured by the OPTECH-RI-V from Union Tool, whose measuring speed is from 300 r/min to 500,000 r/min and resolution is 0.1mm. The supply pressure is 0.5MPa. The spindle temperature is controlled by the circulating cooling water with the inlet temperature of 20 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001042_0959651817698350-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001042_0959651817698350-Figure13-1.png",
        "caption": "Figure 13. Manipulator configuration stage III.",
        "texts": [
            " Tip displacement expressions for stage II can be given as X tip=X CM + rcf + l1cf 1 + l2cf 1 2 + l3cf 1 2 3 + l46cf 1 2 3 4n + l67cf 1 2 3 4n 6n \u00f023\u00de Y tip=Y CM + rsf + l1sf 1 + l2sf 1 2 + l3sf 1 2 3 + l46sf 1 2 3 4n + l67sf 1 2 3 4n 6n \u00f024\u00de where cf =cosf ,cf 1 =cos(f +u 1), ...,cf 1 2 3 4n 6n =cos(f + u 1+u 2+u 3+u 4n+u 6n) and so on. The terms l46 and l67 are link lengths for new configured planar manipulator. It is notable that the superscript \u2018\u2018*\u2019\u2019 with variables denotes the parameter\u2019s values at d1LO . 0, which are known by the sensors used in the model as shown in Figure 3. The joint variables u 4n and u 6n of the new configured robot can be evaluated easily. Stage III. This stage is schematically represented in Figure 13. The circumstance when the manipulator could avoid both the obstacles is shown in Figure 13. Now, to satisfy both the conditions as discussed in stage II, joint 4 is made at FJP and joint 1 is again resumed which was kept at FJP in stage II. Hence, joints 1 and 6 are now active. Now, a problem may arise by selecting the joint 1 again as active (which was active in stage I and then made FJP in stage II) if it causes a collision of the manipulator\u2019s first section with the obstacle 1. To resolve this problem, it must always be assured that d1LO 0 and motion of the manipulator\u2019s first section should be away from that obstacle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003549_j.jmatprotec.2018.08.017-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003549_j.jmatprotec.2018.08.017-Figure7-1.png",
        "caption": "Fig. 7. Finite element mesh generation: (a) GMAW process and (b) VPPA\u2013GMAW process.",
        "texts": [
            " The transient thermal conduction equation is given as: \u239c \u239f \u23a1 \u23a3\u23a2 \u2202 \u2202 + \u2212 \u2202 \u2202 \u23a4 \u23a6\u23a5 = \u2202 \u2202 \u239b \u239d \u2202 \u2202 \u239e \u23a0 + \u2202 \u2202 \u239b \u239d \u2202 \u2202 \u239e \u23a0 + \u2202 \u2202 \u239b \u239d \u2202 \u2202 \u239e \u23a0 +\u03c1C T t v T y x k T x x k T y x k T z q( )P v0 (14) where \u03c1 is the density, Cp is the specific heat capacity, T is the temperature, k is the thermal conductivity, qv is the volumetric heat source term, and v0 is the welding speed. The boundary conditions are as follows: = =t T x y z T0, ( , , , 0) 0 (15) On the workpiece: \u2212 \u2202 \u2202 = \u2212k T z \u03b1 T T( )0 (16) where t is the time, \u03b1 is the combined heat transfer, and T0 is the ambient temperature. Fig. 7. Finite element mesh generation: (a) GMAW process and (b) VPPA\u2013GMAW process. The simulation model was of an Al\u2013Cu\u2013Mg alloy plate with dimensions of 250\u00d7 100\u00d76mm. A non-uniform finite mesh was applied based on the eight-node hexahedral elements. As indicated in Fig. 7, a fine mesh was used in the weld zone and heat-affected zone to ensure enough numerical accuracy, whereas a coarser mesh was used farther from the weld center to reduce the calculation time. The filling process was achieved by a dead-activation element technique in the model. The temperature fields in the VPPA\u2013GMAW process were calculated for different welding conditions, based on the thermal properties of Al\u2013Mg\u2013Cu alloys and the developed heat model. Fig. 8(a)\u2013(d) exhibit the predicted temperature fields of weld cross sections, under the different experimental conditions given in Table 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002523_j.measurement.2016.02.017-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002523_j.measurement.2016.02.017-Figure11-1.png",
        "caption": "Figure 11 Forces and their distances from certain load points on X- axis coordinate",
        "texts": [
            " Afterwards, that point to be considered as a load point in the normal force. Secondly, it should be measured longitudinal and lateral distance of this point until the exact location of the cutting forces at the tip of cutting tool, and distance from the machine bed according to real dimensions of the CNC machine. The values are distances from center of gravity (G). Eq. 4 Eq. 5 Where here; is the total mass of system. The load points of cutting force and the weight force on X-axis are necessary to calculate normal force at contact surfaces. Figure 11 a-c, shows the forces and their distance from the certain points relative to coordinate axes. According to Figure 11 to Figure 13, as a rigid body model, the system of forces can be projected to each point on a body with the corresponding moment. Torque, moment or moment of force, is a force to rotate an object around an axis. Just as a force is a push or a pull, a torque can be thought of as a twist to an object. Mathematically, torque is defined as the cross product of the distance vector and the force vector, which tends to produce rotation. The force vector of cutting force components and weight force creates a torque or couple around the center of gravity. This torque causes reaction forces on the contact surfaces, which is illustrated in Figure 13. The reaction forces increase the friction forces as depicted in Figure 14. The exerted torques and forces can be calculated form the following equations for any component in X, Y and Z directions: From fig. 11 to 13 (a), From fig. 11 to 13 (b), From fig. 11 to 13 (b), From fig. 11 to 13 (c), Because of the forces affecting on the objects in the stationary state. Then: The free body diagram shown in Figure 14. The total friction force in X-axis is as following: Eq. 6 Where; (refer to Figure 8). The static and kinetic coefficient friction ( ) are considered to identify the friction force in X- axis in both of stationary and moving table on the guideways. Eq. 7 Eq. 8 Where, is a total static friction force and is total kinetic friction forces in X-axis linear guideways. The cutting force and normal force result in the friction force on Z direction during cutting operations are shown in Figure 15"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000049_robio.2013.6739433-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000049_robio.2013.6739433-Figure5-1.png",
        "caption": "Fig. 5. Connection diagram of the system.",
        "texts": [
            " While swimming forward, the vehicle can yaw to the left (or right) by setting the vibration amplitude and vibration frequency of the foils on the right side larger (or smaller) than the left. Yawing motion while staying in situ (Fig. 4f). While staying still, the vehicle can yaw to the left (or the right) by setting the offsets of the left pair foils leftX and the right pair foils rightX to  and 0 (or ( 0 and )) respectively. D. Controller design At present study, the vehicle is operated by remote operating handle (F510, Logitech). As shown in Figure 5, an upper monitor (any laptop with the relative software) is used to recognize the operation signals of the handle and generate motion commands to the vehicle. By calculating the navigation information, the monitor also used to display the motion and attitude status of the vehicle in real time. In order to transmit signals between the vehicle and the upper monitor, a CAN-RS232 transducer is also required. As to the functions of the buttons on the operating handle, they are annotated clearly in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002340_1350650114562485-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002340_1350650114562485-Figure1-1.png",
        "caption": "Figure 1. Angular contact ball bearings in divergent \u2018\u2018O\u2019\u2019 and convergent \u2018\u2018X\u2019\u2019 arrangements.1",
        "texts": [
            " Politechniki 6 90-924 Lodz, Lodz 90-924, Poland. Email: jarka@p.lodz.pl at Gebze Yuksek Teknoloji Enstitu on December 19, 2014pij.sagepub.comDownloaded from application, the inner radial clearance in the bearings should be larger than a normal one, so that under the influence of axial force, the operation angle is bigger than zero, similarly to the case of the angular contact ball bearings.1 Preload of angular contact ball bearings refers at the same extent to convergent (\u2018\u2018X\u2019\u2019) and divergent arrangements (\u2018\u2018O\u2019\u2019) (Figure 1), although they are achieved with different technical methods in these arrangements. Preload of bearings results from introducing initial tightness, which is a negative operating clearance. Yet, in bearing arrangements, both the negative and the positive operative clearance occur. For example, operating clearance would be positive, i.e. during work the bearing should demonstrate residual clearance, in bearing systems of vehicle wheels. In many other applications (e.g. bearings for machine tools spindles, pinion bearings in vehicle drives, bearings of small electric engines and oscillation mechanisms), negative clearance, i",
            " Therefore, the influence of the following factors on loading of the balls in the bearing has been considered: \u2013 radial and axial load influencing the bearing shaft, \u2013 elastic deflection of shaft, causing the tilt of inter- nal rings in bearings and \u2013 initial tightness. Figure 2 presents the set of considered loads (example referring to one plane only). The external loads coming from hypothetical gear wheels (Fx, Fy) are the basis for calculating forces Rx i Ry operating in bearings. For calculating the forces, it has been assumed that the reactions of bearings are concentrated in the intersection of the balls operation, i.e. in accordance with the dimension \u2018\u2018c\u2019\u2019 marked in Figure 1, different from the distance \u2018\u2018L\u2019\u2019 between bearings. The problem of determining forces in bearings is statically unassigned. In order to solve it, the authors used a procedure used before,2,8 consisting of iterative searching of such dislocations in bearings (shifts and tilts), for which all conditions of equilibrium are fulfilled: \u2013 compliance of radial reactions of bearings with radial forces influencing the bearings in accordance with the scheme in Figure 2, \u2013 compliance of the sum of the axial reactions of bearings with the sum of the external axial loads, \u2013 compliance of angles of bearings tilts and angles of deflection of shaft under bearings, \u2013 compliance of bending moment in bearings and bending moments considered when calculating the line of shaft deflection and \u2013 compliance of the sum of axial shifts in bearings with the value of initial tightness"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003599_j.procir.2018.04.033-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003599_j.procir.2018.04.033-Figure2-1.png",
        "caption": "Fig. 2 Designs of standard test parts",
        "texts": [
            " This article is mainly based on the existing bibliography and t auth rs\u2019 experience to list the typologies of geometrical defect on the EB pr cess with material Titanium (Ti6Al4V). It could be noticed that the quality of the part made by AM technologies has strong dependency with the actual feature geometrical designs [17]. Therefore, standard test parts with many representative features were proposed by various groups [18, 19, 20, 21]. Illustrations of the most common examples are available in Fig. 2. In general, three main objectives are led with these different tests [19]:  Evaluate the geometric quality of the features produced by machines.  Compare the mechanical properties of features or geometries.  Search for the optimum process parameters for features and geometries. As part of our study, we focus on the first objective: the studies of geometric and dimensional variations. Based on the review of [18, 19, 20, 21], it could be noticed that the tolerances of parts were measured under six types of geometric characteristics, such as straightness, parallelism, perpendicularity, roundness, concentricity and the accuracy of the feature position by Yang et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002953_1045389x16679018-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002953_1045389x16679018-Figure2-1.png",
        "caption": "Figure 2. Piezoelectric plate",
        "texts": [
            " In this case, it is possible to write sk p =Cppek pp +Cpnek pn ekT p Ek sk n =CT pnek pp +Cnnek pn ekT n Ek ~D k = ek pek p + ek nek n + ekEk \u00f024\u00de where ek p = 0 0 0 0 0 0 ek 31 ek 32 ek 36 2 4 3 5ek n = ek 14 ek 15 0 ek 24 ek 25 0 0 0 ek 33 2 64 3 75 \u00f025\u00de The electric field Ek can be derived from the Maxwell equations Ek =DeFk \u00f026\u00de where De = \u2202, x 0 0 0 \u2202, y 0 0 0 \u2202, z 2 4 3 5 \u00f027\u00de and Fk is the electric potential. In this work, two configurations are considered: sensor and actuator. Sensor configuration means that a piezoelectric plate is subjected only to external mechanical loadings, and the resulting deformation state causes the potential distribution. Actuator configuration means that the deformation of the piezoelectric plate is caused by the piezoelectric layers as a consequence of the application of a potential distribution. Both configurations are reported in Figure 2. Further details can be found in Ballhause et al. (2005) and in Carrera et al. (2011a). The potential distribution is a scalar quantity, but for implementation reasons it is convenient to define it as a vector, i.e. Fk = Fk Fk Fk T . In this case, the matrix of differential operators can be expressed as De = \u2202, x 0 0 0 \u2202, y 0 0 0 \u2202, z 2 4 3 5= \u2202, x 0 0 0 \u2202, y 0 0 0 0 2 4 3 5 |fflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl} DeO + 0 0 0 0 0 0 0 0 \u2202, z 2 4 3 5 |fflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflffl} Dez \u00f028\u00de The analysis of a piezoelectric plate can be conducted by means of the PVD (see (11)) extended to the electro-mechanical problem, that states XNL k = 1 Z Vk dekT p sk p + dekT n sk n dEkT ~D k dVk = XNL k = 1 dLk ext \u00f029\u00de Plate geometry is reported in Figure 1, the reference surface is denoted as O and its boundary as G",
            " In the case of an actuator configuration, a potential distribution is applied to the plate, and the value of the potential is set to 1 V at the top and to 0 V at the bottom. The sensor and actuator configurations will be defined as problem 1 and problem 2, respectively. In the case of sensor configuration, the pressure is assumed as in (55), while, in the case of actuator, the potential distribution is assumed as F=F sin mp a x sin np b y \u00f056\u00de where m= n= 1. The reference system layout and the representation of the two configurations are reported in Figure 2. F is set equal to 1. The definition of a BTD is possible if a reference solution is available. The BTDs reported in this work are based on the solution computed by means of the LD4 model that in previous works was demonstrated to be in excellent agreement with the elastic solutions. The interested readers can refer to Carrera (2003) for the mechanical case, to Carrera (2002) for the thermal stress analysis and to Ballhause et al. (2005) for the piezo-mechanic analysis. The results are provided in terms of maximum amplitude of sinusoidal distribution, so the transverse displacement uz and normal stresses sxx, syy and szz are evaluated in the centre of the plate (a=2, b=2), the shear stress sxz in (a, b=2), syz in (a=2, b) and sxy in (a, b)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure2-1.png",
        "caption": "Fig. 2. Non-traditional undercutting \u2014 type II: a) non-undercutting; b) undercutting \u2014 type IIa; c) undercutting \u2014 type IIb.",
        "texts": [
            " 1 PA \u00bc r sin\u03b1 ; PL \u00bc ha\u2212X\u00f0 \u00de= sin\u03b1 , the traditional condition for and ta non-undercutting \u2014 type I finally is written in the following way [2,18,19](2) \u00f02\u00de In the inequality (2) x=X/m is an addendummodification coefficient of the rack-cutter, ha\u204e=ha/m is an addendum coefficient of basic rack tooth, z represents the number of teeth, \u03b1 is the profile angle of the rack-cutter, andm is themodule of the gear (m=2r/z). Minimum addendummodification Xmin of the rack-cutter, eliminating the undercutting \u2014 type I, is defined by the equation Xmin \u00bc xminm \u00bc h a\u22120:5z sin2\u03b1 m : \u00f03\u00de Xmin corresponds to the so-called boundary case \u2014 type I (Fig. 1c), where the tip-line g-g passes through the boundary point A (the point where the line of action contacts the base circle of a radius rb). As already mentioned, the undercutting \u2014 type II is caused by the rack-cutter fillet AF (Fig. 2) in the process of tooth cutting. When this fillet is a circle of a small radius \u03c11 (Fig. 2a) the teeth cut are not undercut. Then the gear fillet fb does not cross the radial line Ob, passing from the center O to the starting point b of the involute profile ba (at X=Xmin point b lies on the base circle of a radius rb). At a comparatively larger radius \u03c12 of the rack-cutter fillet AF (Fig. 2b), an undercutting \u2014 type IIa is obtained, where the fillet gear fb of the generated tooth crosses (cuts) the radial line Ob, but does not cross the involute profile ba. This means that in the presence of undercutting \u2014 type IIa the tooth thickness at the bottom decreases without cutting an involute profile in the vicinity of its starting point b. When the radius \u03c13 (Fig. 2c) of the rack-cutter fillet increases considerably, the fillet gear fq crosses the radial line Ob, as well as the involute profile ba. In this case (undercutting \u2014 type IIb) besides the decrease of the tooth thickness at the bottom, the segment bq of the involute profile is also cut. The essence of undercutting \u2014 type IIa and type IIb is explained on Fig. 3, where undercutting \u2014 type I is avoided by a positive displacement of the rack-cutter at a distance Xmin. At this boundary displacement, the tip-line g-g of the rack-cutter crosses the line of action AB in starting point A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002095_j.jsv.2015.02.010-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002095_j.jsv.2015.02.010-Figure6-1.png",
        "caption": "Fig. 6. Experimental set-up used for measurement of beam vibrations excited by rolling contact [17].",
        "texts": [
            " Nevertheless, to accurately model the viscous damping and its hypothetical nonlinear dependency, the underlying physics \u2013 possibly including dissipation caused by sliding friction, air-squeezing, plastic deformation or material hysteresis \u2013 must be better understood. To that end, the recent article by Winroth et al. [24] is interesting. The proposed methodology is applied to a beam\u2013ball rolling contact for which simulated results are compared to experimental results presented in [17]. In the experimental set-up shown in Fig. 6, a steel ball is allowed to accelerate down a slope and reach a specified velocity when it rolls over a beam with a rough surface. Investigation of beam vibrations is effected for the part between clampings. To study the influence of surface roughness, two beams having different degrees of roughness are considered. The ball roughness rms-level is in [17] confirmed by measurements to be less than 10 percent of that of the smoothest beam. As a consequence, the ball is in the simulations assumed to be perfectly smooth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000599_j.engfailanal.2012.02.008-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000599_j.engfailanal.2012.02.008-Figure15-1.png",
        "caption": "Fig. 15. The field of equivalent stresses on top side of traverse beam.",
        "texts": [
            " 12b), dynamical effects during loading (and unloading) of ladles invoked smaller stress amplitudes (Fig. 12a, approximately 6 MPa). As results from stress charts in Fig. 12 and Fig. 14, the traverse beams are not uniformly loaded and this can be caused by incorrect centering of ladles during their positioning on casting pedestal. On the basis of operator demand that the residual lifespan of supporting structure have to be increased quickly, there was accomplished detailed numerical analysis of influences of proposed modifications to stress state in supporting element of pedestal. In Fig. 15 is given the field of equivalent stresses on upper side of traverse beam of original structure of pedestal for the most danger state of loading. In Fig. 15 is seen locations K and L (see also Fig. 3) of stress concentrations in the corners of hole for the bearing (L) and in location of connection of a beam to the flange of bolted joint (K). Locations K of stress concentration are identical with the locations of crack initiations (see Fig. 3). In locations L (despite high value of equivalent stress) did not arise observable cracks, probably as a result of occurrences of higher radii of rounding (created by welds) as was modeled by the finite element method",
            " 16 is given photograph of upper side of traverse beam after replacement of material in the corner of opening for bearing. In order to decrease stress levels in locations K and L it was recommended to reinforce traverse beams according to Fig. 17, where is given computational model for the finite element method. In Fig. 18 is shown the field of equivalent stresses on reinforced traverse beam for the loading by one full ladle. From comparison of equivalent stress fields on upper side of traverse beam before (Fig. 15) and after (Fig. 18) reinforcement results that reinforcement decreases in locations of stress concentrators K and L stress peaks by approximately 15%. On the basis of analysis obtained during solution problems of loading of casting pedestal the following conclusions can be stated. Due to periodical loading and unloading of wedge joints between traverse beams and middle part the successive drawing out of wedges occurred that results to irregular loading of traverse beam. This problem was solved with protection of upper wedges by stops"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003777_s40799-019-00308-0-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003777_s40799-019-00308-0-Figure1-1.png",
        "caption": "Fig. 1 Application of slewing bearings",
        "texts": [
            "eywords Slewing bearings . Life prediction . Degradation performance description . Health stage division Slewing bearings are key components extensively assembled in wind turbine generators, tunnel-boring machines, tower cranes, military technology and so on. Figure 1 depicts its various applications vividly. For diverse working conditions, there exists several structure forms of slewing bearings including but not limited to (a) single-row ball bearing; (b) double-row ball bearing; (c) cross-roller bearing; (d) three-row roller bearing which are listed in Fig. 2. In addition, it works as a large-size rolling rotational connection between slewing systems with rotational speed ranging from 0.1 rpm to 5 rpm and often bears axial force, radial force and overturning moment under harsh outdoor environments"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001072_s00170-017-0463-2-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001072_s00170-017-0463-2-Figure7-1.png",
        "caption": "Fig. 7 YW7232CNC worm wheel grinding machine",
        "texts": [
            " The last section analyzes the challenge for the new topology modification method, which proves this method could be realized theoretically. In order to strictly prove its functions, two experiments will be conducted. 5.1 Experimental equipment (1) Machine tool This method is based on the external tooth-skipping gear honing technique which uses a worm-shaped honing wheel. The structure of the honing machine is the same as that of the worm wheel grinding machine. So CHMTI YW7232CNC is used in these experiments, as shown in Fig. 7. (2) Measuring instrument The tooth profile deviation and the helix deviation of the processed gear should be measured to provide the results. Klingelnberg P26 Gear Measuring Centre could finish this work very well, as shown in Fig. 8. The accuracy is \u00b11.6 \u03bcm. The experiments mainly certify two things: One is that the new method could accomplish the gear topology modification. The other is that the same honing wheel without any modification could process different modifications using this method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003599_j.procir.2018.04.033-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003599_j.procir.2018.04.033-Figure5-1.png",
        "caption": "Fig. 5 Ideal and Actual Direction of platform motion [23]",
        "texts": [
            " But in this paper, three sources of dimensional inaccuracy could be identified:  The staircase: part orientation and layer thickness have an effect on the staircase error, see Fig. 3. Increasing layer thickness results in more pronounced staircase error.  The error of the position of the energy source: the energy source is imperfectly positioned on the manufacturing platform, see Fig. 4. This figure refers to laser powder bed fusion but we considered that there was the same error with laser beam melting.  The error of platform position, see Fig. 5: there are deviations between actual and ideal motions of manufacturing platform of the machine that introduce errors in the manufactured part\u2019s geometry and dimension. In conclusion, on these defect types based on the bibliography sources, geometry and dimensional variations correspond to the measure of tolerances of the simple feature manufactured by additive manufacturing. The tolerance of part depends on effect of staircase and effect of the machine positional error (source energy and direction of the platform motion)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000484_s11581-014-1068-5-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000484_s11581-014-1068-5-Figure2-1.png",
        "caption": "Fig. 2 Cyclic voltammograms of 0.50 mM EP at the (b) bare CPE, (d) CTPE, (e) DDECPE and (f) DDECTPE. Curves (a) and (c) show the cyclic voltammograms (CVs) of the blank solution at the bare CPE and DDECTPE, respectively. Electrolyte, 0.1 M phosphate buffer solution (pH 7.0); scan rate, 20 mV s\u22121",
        "texts": [
            " 1(A)), the calculated surface concentration is 8.4\u00d710\u221211 mol cm\u22122 for n=2. The electrochemistry of DDE molecule is generally pHdependent. Thus, theelectrochemical behavior of DDECTPE was studied at different pH levels using cyclic voltammetry (Fig. 1(C)). Since one straight line was obtained with a slope value of 54 mV per pH in the pH range of 2.0\u201311.0, there is a transfer of two electrons and two protons in the redox reaction of DDE in the pH range of 2.0\u201311.0 [44]. Electrocatalytic oxidation of epinephrine Figure 2 shows the cyclic voltammetric responses for the electrochemical oxidation of 1 mM EP at DDECTPE (curve f), DDE-modified CPE (DDECPE) (curve e), TiO2 nanoparticle CPE (CTPE) (curve d) and unmodified CPE (curve b). DDECTPE, in 0.1 M phosphate buffer (pH 7.0), showed a well-defined redox reaction (curve c). The anodic peak potential for EP oxidation at DDECTPE (curve f) and DDECPE (curve e) was about 205 and 215 mV, respectively, while at the CTPE (curve d), it was about 390 mV. At the unmodified CPE, the peak potential of EP was about 395mV (curve b)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003234_1350650117738395-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003234_1350650117738395-Figure3-1.png",
        "caption": "Figure 3. -Comp test rig.",
        "texts": [
            " To verify that 1000 cycles per speed step was sufficient and to evaluate the spread between different synchronizers tested under the same conditions, 11 synchronizers were tested at 2500N and 0.42 kg m2. In this test the speed to synchronize was also increased every 1000 cycles until failure. Additionally, one endurance test is included in this paper, which shows a synchronizer completing 200,000 cycles at 2500N, 0.42kg m2, and 920 r/min, which is high where most synchronizers fail. Physical tests were run in a -comp test rig,24 which is schematically shown in Figure 3. The latch cone is connected to the rigs shaft. The inner cone is fixed to ground. The electric motor rotates the flywheel to the set point speed, and the synchronizer is used to brake the flywheel to standstill. A hydraulic actuator is used to generate the shift force. The rotational speed, the applied force, and the resulting cone torque were measured at 1000Hz. The accuracy of the sensors is presented in a previous investigation.23 The coefficient of friction can be determined by rearranging equation (1) to C \u00bc MC sin FAx RC \u00f08\u00de Additionally, a thermocouple (type K) was attached to the inside if the inner cone (6mm from the contact surface) solely to verify that the start of cycle temperature was equal for the different load cases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002969_s00170-017-0213-5-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002969_s00170-017-0213-5-Figure2-1.png",
        "caption": "Fig. 2 Definitions of geometric parameters",
        "texts": [
            " The objectives of this study is to analyze and control the flow field in the SLM chamber of the laminated manufacturing system, for the purpose of capturing all the metal powders in a blow-to-suction device. We conducted numerical simulations across the chamber as well as verified the simulation results experimentally using flow measurement and flow visualization on a 1:1 lab-scale model. The experiment was conducted in a blow-to-suction working chamber, which consists of a trapezoid push nozzle, a cuboid chamber, and a suction device. As shown in Fig. 2, it was built to simulate the industrial 250 mm \u00d7 250 mm SLM chamber. The experimental working chamber was made of acrylic and assembled by the aluminum extrusion tooling. The transverse length and width in z and x directions are 605 and 325 mm, respectively. The height of the working plane can be adjusted between 414 and 464 mm using a pressure-regulating valve with no leakage. The trapezoid push nozzle was designed to have five vertical plates properly arranged and one flow field development area to reduce the turbulence intensity and increase the flow uniformity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure9-1.png",
        "caption": "Figure 9: z-displacement of the flux of force adapted lattice structure",
        "texts": [
            " The course of the lattice structure inside the design space (see figure 3) has been adapted to the internal flux of force for the given constraints and loads (see figure 8). In order to keep the results comparable, the struts\u2019 diameters have been chosen to 2 mm as in case of the periodic structure. So, the course of the structure is the only differing variable in contrast to the periodic structure. The resulting structure has a mass of 83.57 g, which is 43 % less than in case of the periodic structure. Stiffness. Figure 9 shows the simulation result for the z-displacement of the structure. In order to make the results more comparable, the colour scale is the same as in figure 5. The displacement in the point of force application is 0.139 mm, which means a slight improvement compared to the periodic structure. Together with the applied force of 300 N, a stiffness of 2166 N/mm results for the structure. Due to the extensive reduction of the structure\u2019s mass, the ratio of stiffness to mass improves for 81 % towards a value of 25"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001072_s00170-017-0463-2-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001072_s00170-017-0463-2-Figure8-1.png",
        "caption": "Fig. 8 P26 Gear Measuring Centre",
        "texts": [
            "1 Experimental equipment (1) Machine tool This method is based on the external tooth-skipping gear honing technique which uses a worm-shaped honing wheel. The structure of the honing machine is the same as that of the worm wheel grinding machine. So CHMTI YW7232CNC is used in these experiments, as shown in Fig. 7. (2) Measuring instrument The tooth profile deviation and the helix deviation of the processed gear should be measured to provide the results. Klingelnberg P26 Gear Measuring Centre could finish this work very well, as shown in Fig. 8. The accuracy is \u00b11.6 \u03bcm. The experiments mainly certify two things: One is that the new method could accomplish the gear topology modification. The other is that the same honing wheel without any modification could process different modifications using this method. The parameters of the wheel and the gear used in the experiments are shown in Table 1. In the second experiment, the modification parameters are changed. In both experiments, only one honing wheel is employed. The parameters of the gear in this experiment and the modification parameters are in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001908_1.4921493-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001908_1.4921493-Figure1-1.png",
        "caption": "FIG. 1. Schematic representation of powder deposition and application of boundary conditions.",
        "texts": [
            " The simulation model neglects the in-flight heating of the powder particles and assumes that the mass of powder particles is spread with a uniform thickness on the substrate or an already deposited layer. The model also neglects heat loss due to bouncing-off of hot powder particles from the deposit surface. These assumptions are undertaken to keep the model tractable and would perhaps have little effect because of high power density, and rapid melting and solidification of a small pool under the focused laser beam. Figure 1 depicts the solution domain that schematically shows a substrate of size 25 10 2 mm (length width height) and a build of 8 0.14 mm (length width). The solution domain is discretized using 3D eight node element (DC3D8 in ABAQUS) with the temperature as the degree of freedom. The deposition of the powder materials is modeled using activation of a set of discrete elements in each time-step during which the laser is presumed to move a length as the beam radius in the direction of deposition. The transparent elements in Fig. 1 conform to the deactivated elements depicting the powder particles yet to be deposited while the orange, green, and red colored elements refer, respectively, to the substrate, already deposited powder, and the activated, i.e., deposited powder in the current time-step. A nonuniform grid is used with finer elements for the build geometry and relatively coarser elements for the substrate. The minimum and the maximum sizes of the elements considered were 0.1 and 0.8 mm, respectively, with the total number of elements was 94 000"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure4-1.png",
        "caption": "Fig. 4. Common moving frames and coordinate systems considering errors.",
        "texts": [
            " (14)\u2013(16), (20), {Op; \u03b11\u03b12\u03b13} in S1 can be represented by: \u03b11 \u00bc B \u03c61\u00f0 \u00dee p\u00f0 \u00de 1 \u00bc \u03c61 sin 2\u03b4e \u0394\u03bb1\u00f0 \u00de\u2212 cos \u03b4 cos \u03b4e1 \u0394\u03bb1\u00f0 \u00de \u00fe sin \u03b4k1\u00f0 \u00deffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi cos2\u03b4\u00fe \u03c61 sin 2\u03b4 2q \u03b12 \u00bc B \u03c61\u00f0 \u00dee p\u00f0 \u00de 2 \u00bc sin \u03b4e1 \u0394\u03bb1\u00f0 \u00de\u2212 cos \u03b4k1 \u03b13 \u00bc B \u03c61\u00f0 \u00dee p\u00f0 \u00de 3 \u00bc cos \u03b4e \u0394\u03bb1\u00f0 \u00de \u00fe \u03c61 sin 2\u03b4 cos \u03b4e1 \u0394\u03bb1\u00f0 \u00de \u00fe sin \u03b4k1\u00f0 \u00deffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi cos2\u03b4\u00fe \u03c61 sin 2\u03b4 2q 8>>>>< >>>>: \u00f021\u00de According to the coordinate transformation rules from S1 to S2, this common moving frame can also be denoted as: \u03b11 \u00bc \u2212B \u03c62\u00f0 \u00dee g\u00f0 \u00de 1 \u00bc \u2212 \u03c62 cos 2\u03b4e \u0394\u03bb2\u00f0 \u00de \u00fe sin \u03b4 sin \u03b4e1 \u0394\u03bb2\u00f0 \u00de \u00fe cos \u03b4k2\u00f0 \u00deffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi sin2\u03b4\u00fe \u03c62 cos 2\u03b4 2q \u03b12 \u00bc \u2212B \u03c62\u00f0 \u00dee g\u00f0 \u00de 2 \u00bc cos \u03b4e1 \u0394\u03bb2\u00f0 \u00de\u2212 sin \u03b4k2 \u03b13 \u00bc \u2212B \u03c62\u00f0 \u00dee g\u00f0 \u00de 3 \u00bc sin \u03b4e \u0394\u03bb2\u00f0 \u00de\u2212\u03c62 cos 2\u03b4 sin \u03b4e1 \u0394\u03bb2\u00f0 \u00de \u00fe cos \u03b4k2\u00f0 \u00deffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi sin2\u03b4\u00fe \u03c62 cos 2\u03b4 2q 8>>>>< >>>>: \u00f022\u00de In order to reduce the influence of the shift of bearing contact caused by misalignment, the local conjugate theory of gearing is used to design pinion-gear tooth profiles of HNCGs, and the line contact of teeth at every instant is altered into the point contact. However, the conjugation properties of teeth can still be investigated in the normal section plane to tooth surface, and the moving frame of pinion tooth surface can be used as a common moving frame of conjugated surfaces at the contact position. Here, a hypothesis is made that the left tooth flank of pinion contacts with the right tooth flank of gear when the conjugation of teeth is observed from toe to heel. As shown in Fig. 4, according to the geometrical characteristic of tooth surface, at the predesigned contact point P of tooth pair, the common moving frame {P; \u03b1 s1\u03b1 s2\u03b1 s3} for conjugated tooth surfaces can be represented by: where \u03b1s1 \u00bc \u03b11\u03b1s2 \u00bc \u2212\u03b12 sin \u03b1n \u00fe \u03b13 cos\u03b1n\u03b1s3 \u00bc \u2212\u03b12 cos \u03b1n\u2212\u03b13 sin \u03b1n 8< : \u00f023\u00de Actually, \u03b1s1, \u03b1 s2 are unit vectors of principle directions of pinion tooth surface. Thereinto, \u03b1 s1 is a unit vector along the toothlength direction (TL direction) of pinion, i.e. from the toe to the heel, \u03b1 s2 is a unit vector along the tooth-height direction (TH direction) of pinion, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002668_1350650115619610-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002668_1350650115619610-Figure6-1.png",
        "caption": "Figure 6. Finite element mesh of thrust bearing with elliptical shape recess.",
        "texts": [],
        "surrounding_texts": [
            "Pocket pressure poc\nThe variation of pocket pressure for Newtonian lubricant by considering thrust pad rigid and flexible is shown in Figure 12(a). From the result, it is observed that pocket pressure monotonically decreases with increasing value of the tilt parameter. The result shows that deformation of thrust pad results a decrease in the value of pocket pressure. For the selected value of tilt parameter l \u00bc 0:4 and Newtonian lubricant, the change in the value of poc due to consideration of deformation in thrust pad is 1.51%, 1.60%, 1.53%, and 1.65% for circular, elliptical, square, and rectangular shape recess, respectively. This decrease in the value of pocket pressure is due to increase in the value of fluid-film thickness. The pocket pressure of rectangular shape recess has the maximum affect by deformation. Figure 12(b) presents the variation of pocket pressure with respect to the restrictor design parameter for the selected value of tilt parameter. The circular shape recess has the maximum value pocket pressure, whereas the rectangular shape recess has the minimum value of fluid-film pressure. The value of pocket pressure increases with an increase in the value of restrictor design parameter. Figure 12(c) presents the variation of\npocket pressure of thrust bearing operating with psuedoplastic lubricant as a function of tilt parameter l\u00f0 \u00de. For the selected value of tilt parameterl \u00bc 0:4 and psuedoplastic lubricant, the change in the value of poc due to consideration of deformation in thrust pad is 2.06%, 2.24%, 2.12%, and 2.30% for circular, elliptical, square, and rectangular shape recess respectively. From the results of Figure 12, following foremost pattern may be drawn\npoc Tilt 5 poc Parallel\npoc Rigid 4 poc Flexible\npoc Circular 4 poc Square 4 poc Elliptical 4 poc Rectangular\npoc Non Newtonian 5 poc Newtonian\nLubricant flow rate Q\nFigure 13 depicts the variation in the value of lubricant flow rate Q with respect to tilt parameter l\u00f0 \u00de and restrictor design parameter Cs2 for the elastohydrostatic Cd \u00bc 0:5 and hydrostatic Cd \u00bc 0:0\n. The consideration of deformation of thrust pad in the analysis results in a significant reduction in the value of\nat UNIV CALIFORNIA SAN DIEGO on December 12, 2015pij.sagepub.comDownloaded from",
            "Q. The value of lubricant flow monotonically increases with an increase in the tilt parameter. This is due to the fact that for a tilted pad thrust bearing, the requirement of flow is higher to operate bearing. This increase is due to change in the value of fluid-film thickness on the pocket of bearing. The lubricant flow rate of thrust bearing having rectangular recess shape is the maximum and lubricant flow rate of thrust bearing having circular shape recess is the minimum. For the selected value of tilt parameterl \u00bc 0:4 and Newtonian lubricant, the change in the value of Q due to consideration of deformation in thrust pad is 8.96%, 7.95%, 8.38%, and 7.84% for circular, elliptical, square, and rectangular shape recess respectively. As seen in Figure 13(b), the value of lubricant flow rate continuously increases with an increase in the value of restrictor design parameter. The variation of lubricant flow rate for pseudoplastic lubricant as a function of tilt parameter is depicted in Figure 13(c). For the selected value of tilt parameterl \u00bc 0:4 and pseudoplastic lubricant, the change in the value of Q\ndue to consideration of deformation in thrust pad is 8.03%, 6.23%, 7.36%, and 6.08% for circular, elliptical, square, and rectangular shape recess, respectively. Following important pattern is observed in the value of lubricant flow rate\nQ Tilt 4Q Parallel\nQ Rigid 5Q Flexible\nQ Circular 5Q Square 5Q Elliptical 5Q Rectangular\nQ Non Newtonian 4Q Newtonian\nLoad-carrying capacity F0\nAs seen in Figure 14, the load-carrying capacity is a strong nonlinear function of tilt parameter. The loadcarrying capacity of thrust bearing having square shape recess is maximum. Maximum influence of tilt\nat UNIV CALIFORNIA SAN DIEGO on December 12, 2015pij.sagepub.comDownloaded from",
            "at UNIV CALIFORNIA SAN DIEGO on December 12, 2015pij.sagepub.comDownloaded from"
        ]
    },
    {
        "image_filename": "designv11_13_0002645_978-3-319-26687-9_7-Figure7.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002645_978-3-319-26687-9_7-Figure7.1-1.png",
        "caption": "Fig. 7.1 Tracking error coordinates. The bold triangle and the light rectangle indicate the real UAV and the VTV, respectively. The bold dash line indicates the desired path",
        "texts": [
            "2m/s2, u2max = 28deg/s. Following the approach developed in [33], this section introduces a virtual target vehicle (VTV) that is constrained to move along a given desired geometric path (xd(\u03b3 ), yd(\u03b3 )) parameterized by the variable \u03b3 \u2208 R, and defines the error vector between the position of the VTV and the actual UAV through the use of a Serret\u2013 Frenet frame, which can be viewed as the body frame of the VTV that moves according to a \u201cconvenient velocity\u201d, effectively yielding an extra control input. Figure7.1 captures the setup, where the dashed line denotes the desired path, and the light rectangle represents the VTV. The desired course heading \u03c7d and the curvature \u03c3d are related by x\u2032 d = cos\u03c7d y\u2032 d = sin \u03c7d \u03c7 \u2032 d = \u03c3d (7.3) where the prime symbol denotes the first derivative of a variable with respect to the parameterization variable \u03b3 . Note that one way to obtain the desired geometric path (xd(\u03b3 ), yd(\u03b3 )) is to solve (7.3) with a given initial condition (xd(0), yd(0), \u03c7d(0)) and specified input signal \u03c3d(\u03b3 )"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000950_6.2017-0011-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000950_6.2017-0011-Figure2-1.png",
        "caption": "Figure 2. Unmodified E-Flite AS3Xtra aircraft (left) and modified aircraft showing wing in forward-swept configuration (right).",
        "texts": [
            " Section V develops a nonlinear flight dynamics model using the aerodynamic coefficients collected in the wind tunnel. The model is then used in combination with numerical trajectory optimization methods to calculate and simulate perching trajectories (Section VI). Theses numerical experiments show that the forward-swept wing is capable of achieving \u201cmore optimal\u201d perching trajectories that require less actuator effort. II. Variable-Sweep Wing Design The starting point for our wing-morphing aircraft was the low-cost flat-foam hobby aircraft shown in Figure 2, which has a wingspan of 42 centimeters and a mass of 37 grams. Several objectives guided the design of the wing sweeping mechanism. First, we sought to make minimal modifications to the original airplane. Second, the mechanism needed to be extremely lightweight to minimize the impact on flight performance. Lastly, we wanted to drive the mechanism with the same small linear servos used to actuate the control surfaces on the airplane. Figure 3 depicts our single-degree-of-freedom modified sliding-crank mechanism design. The red links are attached to the wings, while the yellow link is fixed to the fuselage. The pivot point locations were chosen to provide a 25\u00b0 sweep angle and to allow the ailerons to clear the fuselage. Several prototypes have been constructed from lasercut fiberglass and carbon fiber to provide high stiffness and low weight. A fully assembled airplane with the sweeping mechanism installed is shown in Figure 2. In the remainder of the paper, we analyze the aerodynamic properties of the airplane at the extreme configurations of 0\u00b0 and 25\u00b0 and leave flight experiments with the variable-sweep aircraft to future work. To gain insight into the aerodynamics of the forward-swept wing configuration, experiments were conducted at Harvard University\u2019s Concord Field Station (CFS) wind tunnel. Measurements of aerodynamic lift, drag, and moment were made over a range of angles of attack from \u221225\u00b0 to +75\u00b0. In addition, the roll moment generated by an aileron deflection of 20\u00b0 was measured over the same range of angles of attack to identify differences in high-angle-of-attack control authority",
            " At fast and slow speeds, the duration of flights that the birds were willing to sustain was less than this. The maximum speed of each bird was defined as the highest speed at which the bird was willing and able to maintain position in the wind tunnel for 15 s. Doves learned to fly steadily more slowly, requiring as much as 1 month of training to achieve flights of more than 2 min duration at any given speed. Design of the wind tunnel The Harvard-Concord Field Station (Harvard-CFS) wind tunnel is an open-circuit tunnel with a closed jet in the flight chamber, designed and constructed in 1998\u20131999 (Fig. 2). It has a working section 1.2 m\u00d71.2 m in cross section and 1.4 m in length and can operate at wind speeds from 0 to 28.5 m s\u20131. Air is moved through the tunnel by a 55.9 kW (75 horsepower) direct current motor (General Electric, Inc.) and 1.4 m diameter fan assembly (AFS-1.4 Series axial flow fan, SMJ Inc.) equipped with a built-in silencer (1.4 LCP series) to reduce noise levels. Barlow et al. (1999) and other sources cited therein were used to design the tunnel. Air is first pulled through a settling section, measuring Figure 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003437_s12046-018-0837-7-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003437_s12046-018-0837-7-Figure1-1.png",
        "caption": "Figure 1. Benchmark component.",
        "texts": [
            " Further in this research work, experimental investigations had been carried out by varying two controllable factors of the IC process, namely drying time of primary coating and mould thickness on various properties such as surface roughness, dimensional accuracy and hardness of the castings prepared by this combined route. Also, the biocompatibility of the casted implants prepared by this route (FDM ? CVS ? SM ? IC) had been evaluated by invitro testing. In this research work, the most commonly used biomedical implant (hip joint) as shown in figure 1 has been selected as the benchmark component. The CAD model of the component had been made using CREO 2.0 software and then converted into .STL file. The .STL file serves as an input to the FDM machine to fabricate 3D model of acrylonitrile butadiene styrene (ABS) material followed by the CVS process to enhance the surface finish. It has been observed that CVS process dramatically improves the surface finish of ABS replicas. The results were in line with the observations made by other investigators [19, 26]",
            " Control log of experimentation along with the results of surface finish, dimensional accuracy and hardness of the castings have been shown in table 2. In this work, surface roughness (Ra) and dimensional deviation (Dd) has been taken to represent the surface finish and dimensional accuracy, respectively. Surface roughness (Ra value) of all the castings has been measured at femur part by Mitutoyo SJ-210 roughness tester at 0.5 mm/s stylus speed and cutoff length at 0.25 mm. One radial dimension (as shown in figure 1) of all the PU replicas and corresponding castings was measured by coordinate measuring machine (CMM) and deviation has been measured. The surface hardness of the casted hip joints was measured by using Vickers hardness tester (HVS-1000BVM) as per ASTME-384 standard. In order to minimize the variation, three readings were taken for each response and the average of the three values was considered as the final output. Figure 4 shows the casted components and figures 5\u20137 indicate the variation of output response with the experiment number"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000368_aero.2013.6496903-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000368_aero.2013.6496903-Figure1-1.png",
        "caption": "Fig. 1. Typical Unguarded Wiring Configuration",
        "texts": [
            " Cable lengths in Zone C could range from 0 to 1000 feet. In general, any electronics such as a preamplifier or signal conditioner would be located within Zone C to take advantage of the lower temperatures. Electronics would be located within Zone B only if absolutely necessary and rarely in Zone A. II. HARDWARE IMPLEMENTATIONS A. Probes and Cables Two types of probes and two cable configurations are typically use for capacitive sensors. Probes can be unguarded or guarded, and cables can be coaxial or triaxial. Unguarded probes (see Figure 1) are the simplest, consisting of a center sensing tip surrounded by grounded housing. The tip is insulated from the outer housing similar to the center conductor of a spark plug. Most sensors have circular symmetry, although other shapes such as stripes can be used. The circular shapes are more common in part because high temperature components can be manufactured more easily compared to more-complex shapes. Furthermore, if the probe housing is threaded, it can be positioned at a pre-determined tip-to-blade gap without concern of rotational alignment to the blade",
            " Guarded probe configurations generally result in a wider measurement range compared to unguarded configurations. However, as seen below, use with thin blades affects the capacitance vs. gap relationship and can cause difficulties for such linearized systems. In addition, linearized systems typically have much shorter gap ranges compared to systems whose output is proportional to capacitance. Coaxial cables, which consist of a center conductor surrounded by a shield, are generally used for unguarded probes. As seen in Figure 1, the center conductor of the cable connects the preamplifier output to the probe's sensing tip. The outer shield is usually connected to the engine case at one end and to the preamplifier ground at the other end of the cable. Triaxial cables (see Figure 2) have an additional insulated shield between the center conductor and the outer shield. The center conductor is connected between the preamplifier and the probe's sensing tip. The inner shield of the cable is connected to the guard electrode at the probe and to a buffer amplifier at the preamp"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002557_icelmach.2014.6960311-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002557_icelmach.2014.6960311-Figure15-1.png",
        "caption": "Fig. 15. The geometry of the 4-pole, 24-slot hysteresis motor",
        "texts": [
            " Based on the Preisach density distribution identified from the constructed major and inner hysteresis loops of Fig. 9, the major and inner hysteresis loops computed from Preisach model are shown in Fig. 13. Fig. 14 shows the derived first order transition curves. -0.8 -0.4 0 0.4 0.8 -1600 -800 0 800 1600 B ( T ) H (A/m) Fig. 14. Preisach model computed first order transition curves based on the identified parameters Another application example is a 4-pole, 24-slots, 550W, 1500rpm hysteresis motor, as shown in Fig. 15. The laminated stator core is set with normal BH curve, and a hysteresis material is used for the solid rotor core. The motor is operated at locked rotor with balanced three-phase AC voltage supply. The major hysteresis loop constructed based on the proposed algorithm is used to identify the parameter of the static play model [11] for the rotor hysteresis material. The simulated three-phase current waveforms are shown in Fig. 16. In Fig. 17, the total hysteresis loss is compared with the power calculated by multiplying torque with the synchronous speed, which equals to the electro-magnetic power delivered from the stator to the rotor via air gap when the negative-sequence current component decays to zero at the steady state"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001706_s11071-018-4696-x-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001706_s11071-018-4696-x-Figure3-1.png",
        "caption": "Fig. 3 Inverted pendulum",
        "texts": [
            " Then, the discrete-time PWA systems can be obtained by discretization. Using Theorem 1 in Sect. 3, we can obtain the feedback gains K i and mi of the each subsystem and the domain of attraction is estimated. The maximal domain of attraction can be obtained by solving optimization problem2.Wecompare the simulation results for \u03c3 having different values. The update rate is defined as the ratio of the number of events that occur and the total number of measurements. Example 1 Consider the inverted pendulum system as shown in Fig. 3 which is the actual model of inverter pendulum. \u03b8 indicates angular displacement of pendu- lum. Let ( x1 x2 )T = ( \u03b8 \u03b8\u0307 )T . The inverter pendulum model with saturation level u0 = 5 is considered as [34]:{ x\u03071 = x2 x\u03072 = \u22120.1x2 + sin(x1) + sat(u) The characteristic of nonlinear function sin(x1) is modeled by PWA function. The approximation effect is shown in Fig. 4. The approximation of the equation is described by sin(x1) = \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u2212 0.85x1 \u2212 2.6, \u2212 4 < x1 < \u22122 \u2212 0.9, \u2212 2 < x1 < \u22121 0.9x1, \u2212 1 < x1 < 1 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003816_s12541-019-00014-2-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003816_s12541-019-00014-2-Figure5-1.png",
        "caption": "Fig. 5 Three-dimensional model of the blade bearing of the CPP",
        "texts": [
            " However, the values of the components of the torsor are verified before each simulation to ensure that all of them satisfy the constraint conditions. Figure\u00a04 displays the flow diagram of the modified Unified Jacobian\u2013Torsor model for the statistical tolerance analysis based on Monte Carlo simulation. 4.2 Finite Element Model of\u00a0the\u00a0Blade Bearing of\u00a0CPP After obtaining the assembly deviation caused by the dimensional and geometric tolerances, the next step is to derive the deformation caused by the centrifugal load and temperature. As shown in Fig.\u00a05, the 3D (three dimensional) model of the blade bearing is axis symmetric. To reduce the computational time, only a quarter is constructed, and its material properties are the same as in our previous study [32]. Before the FEA, to integrate the deformation into the assembly deviation, the mating surfaces of each part of the blade bearing should be discrete and it should be ensured that the each pair of nodes on the upper and lower contact surfaces have a one-toone correspondence. Figure\u00a06 exhibits the discretization method for two different planes such as a rounded plane or square plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002182_btpr.2010-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002182_btpr.2010-Figure3-1.png",
        "caption": "Figure 3. Schematic representing of the process of coimmobilization of the two enzymes on silica microsphere.",
        "texts": [
            " Figure 2 shows CD spectra of the free enzymes and immobilized enzymes. The amount of free enzymes was equivalent to that of enzymes immobilized on MWNTs. In the measurement of the CD spectrum of the conjugates, the CD spectrum of MWNTs was recorded as a control. The secondary structures of the enzymes appear to have been reduced after immobilization. Characterization of coimmobilized a-amylase and glucoamylase The process of coimmobilization of the two enzymes consisted of two steps as schematically illustrated in Figure 3. The coimmobilization of the enzymes has been confirmed by the SEM images (Figure 4). Silica microsphere itself presents a smooth surface with a diameter of about 7 mm (Figure 4a). In the first step, CNT-immobilized glucoamylase (CNT-GluA) was adsorbed on silica microsphere (Figure 4b), CNT-GluA is clearly observed on the surface of the microsphere (Figure 4b). In the second step, the silica microsphere/CNT-GluA was coated with a sol-gel layer, in which CNT-immobilized a-amylase (CNT-AA) was entrapped, and the surface of the microsphere becomes smooth (Figure 4c)",
            " a-Amylase and glucoamylase with a weight ratio of 1:1 were encapsulated by a sol-gel layer, which coated the outside of the silica microspheres. Such coimmobilized enzymes were used to hydrolyze starch. It was demonstrated that the catalytic activity of the coimmobilized enzymes retained 78 6 3% of the activity of the mixture of free enzymes. Compared to the random coimmobilization on carbon nanotubes, the activity of the coimmobilized enzymes has been improved. Possibly it is because that the encapsulated enzymes were in a desired ratio of 1:1. In the sequential coimmobilization, as illustrated in Figure 3, the ratio of coimmobilized enzymes is controlled in an easy way, as the amount of each enzyme which is individually immobilized on CNTs can be determined accurately. In terms of the catalytic activity, the sequential coimmobilization exhibited advantages over the encapsulation coimmobilization. The sequentially coimmobilized enzymes exhibit a higher relative activity (95%) than the encapsulated enzymes. In the sequential coimmobilization, carbon nanotubes, in addition to as supports for the enzymes, can function as scaffolding within the matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000011_carpi.2012.6473371-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000011_carpi.2012.6473371-Figure6-1.png",
        "caption": "Figure 6 shows a generic sketch of a petrochemical processing facility. The petrochemical distillation process commonly starts with crude oil being pumped from large tanks through a furnace system to vaporize the crude. Subsequently, it is split into various valuable products (e.g. petroleum, kerosene etc.) in refining reactors or columns. The corresponding structures are commonly large enough to be interesting for aerial inspection applications. Accordingly, relevant components are listed and inspection applications are proposed hereafter.",
        "texts": [],
        "surrounding_texts": [
            "The overall working principles of the various existing types of thermal power plants are essentially quite similar. In general heat is produced either through burning of fuels (oil, coal or gas) or through nuclear fission. The produced heat is used to generate steam from water. The resulting steam is fed to steam turbines driving generators finally producing the desired electricity. Figure 2 presents a sketch of a typical fuel based thermal power plant. Three main structures which are of particular interest for aerial inspection have been identified in this study. Specifications and possible inspection tasks for each of these facility structures are listed next. 1) Cooling Tower Cooling towers are commonly cleaned, inspected and maintained every 10 to 12 years. The corresponding inspection tasks focus on checking structural integrity and are mostly visual. The respective structures are commonly made from steel reinforced concrete, can be up to 200 meters high and up to 50 meters in diameter (Figure 3). Due to the vast size, elevator systems, cranes or climbing utilities must be installed to search for structural damages and reinspect previously repaired areas. To increase inspection efficiency and avoid endangering inspection personnel, aerial inspection robots are thus considered a viable alternative to the more laborious, classical inspection methods. Expected inspection tasks for aerial robots are autonomous scanning of the outer surface of cooling towers while the tower may be on- or offline and scanning of the interior surface during shutdown periods. Generating at least crude 3D representations by mapping the structure is also desirable. Additionally, the aerial inspection vehicle may be required to carry and attach supporting devices (pulley systems or hooks) to specific parts of the tower to pull heavier cleaning machines to different points on the structure. Particular challenges for aerial inspection may be environmental disturbances such as wind gusts while scanning along the outer surfaces of the cooling tower or the elevated humidity and temperature as well a vertical convection winds when flying inside the tower. 2) Boiler system Boilers are large, rectangular chambers made from carbon steel tubes within which water or steam is transported. Through injection nozzles fuel is injected and burned in the boiler furnace, generating the required heat to transform the water in the boiler tubes into steam (Figure 4). Average boilers may be 15 to 35 meters wide and 50 meters high. In general, boiler systems are inspected and maintained in periods of 5 to 6 years. Inspection is usually focusing on corrosion on the tubes and crack detection in particular on tube welds. Moreover, condition assessment of the injection nozzles and the super- and reheater pipework is required. The inspection process commonly starts with a visual preinspection through small windows embedded in the tube walls allowing a limited view on the boiler interior. Often scaffolds are installed next to allow a detailed visual inspection. Subsequently, a cleaning process is started to remove combustion deposits from the boiler walls. Cleaning is also necessary to prepare the structures of interest for NDT measurements e.g. using ultrasonic testing (UT). Damaged nozzles and pipes are then cut from the boiler structure and replaced with new components. Note that access to the boiler interior is generally only possible through small manholes on different levels of the structure. Also note that in the early phases of the boiler maintenance process, no light source of any kind is available inside the furnace chamber. Aerial inspection robots are expected to simplify and extend the visual assessment of the boiler walls, injection nozzles and the super- and reheater pipework at the top of a boiler. All structures must be visually scanned possibly also generating at least crude 3D maps of the boiler chamber. Attaching supporting devices (e.g. using magnetic attachment) to expedite installation of scaffolds or heavy cleaning machinery is desirable as well. Finally, remote inspection by contact using advanced NDT tools is the ultimate goal for aerial inspection robots within this type of plant infrastructure. The third type of structure to be inspected by aerial inspection robots are the chimney stacks commonly found in thermal power plants (Figure 5). Depending on the employed fuel the chimney layout varies. Coal fired power plants usually require large scale chimneys, whereas gas fired power plants have relatively small chimneys. The corresponding diameters range from 0.5 to 2 meters and heights may vary from 20 up to 200 meters. Commonly found materials are concrete and brick and thus structural defects and the condition of previous repairs are of interest. For aerial inspection visually scanning and mapping the chimney exterior and if possible the chimney interior (flying into the chimney) is of interest. For smaller chimneys safely landing on top of the chimney and lowering e.g. a miniature camera system into the chimney structure may also be desired. Besides natural wind gusts, additional convection winds produced by the chimney structure must be expected. As for the boiler, light is limited inside the chimney. 1) Flare Systems Flare systems such as in Figure 7 serve as safety backups in refineries. In case of problems in a process unit, excess gases are combusted in a controlled manner. Heights of typical flare systems may vary between 70 to 90 meters. In general visual inspections are performed when the processing facility is shut down and the flare can be safely turned off. For aerial inspection visual scanning of flares may be desired during shutdown periods or on while the plant is online. In these cases the heat emitted by the flares may prohibit close proximity flights. Monitoring the temperatures on the UAV might also be required to avoid excessive heat exposure. Petrochemical furnaces are used to heat or vaporize crude oil or other intermediate products during the refining process. They are usually fired with gas or crude from the bottom of the furnace barrel (Figure 8). Inspections are planned every 6 years and depending on the fuel (gas or crude) cleaning may be required in periods of 2 to 6 or 10 years. Diameters are commonly between 4 to 10 meters and thus flying inside the structure is of interest. Similar to boilers visual inspection as well as UT measurements of the furnace coils is required. 3) Chimney stack Chimney stacks in petrochemical processing industries are inspected on a regular basis every 5 to 10 years. Their working principle and dimensions are commonly quite similar to chimney stacks in thermal power plants. Possible application scenarios for aerial inspection robots are thus comparable to the scenarios described for power plant chimneys. Tanks represent vital buffering systems for crude oil and distillation products in a petrochemical processing facility. Since the throughput of a refinery is usually exceeding the tanks storage capacity by factors, it is essential to ensure that all tanks are operational at all times. Accordingly, proper lifetime assessment is crucial to avoid problems with unexpected tank failures possibly requiring repairs outside of planned servicing periods. They are inspected on a 5 year basis for wall thinning due to corrosion. Visual inspections are mainly of interest to efficiently plan the detailed NDT inspection. Diameters of tanks as depicted in Figure 9 may vary from 5 to 40 meters. Accordingly, interior or exterior aerial inspections are of interest."
        ]
    },
    {
        "image_filename": "designv11_13_0002832_978-981-10-2374-3_6-Figure6.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002832_978-981-10-2374-3_6-Figure6.1-1.png",
        "caption": "Fig. 6.1 Chattering phenomenon",
        "texts": [
            " Hence, the classic sliding mode control that satisfies the above condition is given by Riachy (2008): u = \u2212 ( \u2202s \u2202x b (x) )\u22121 ( \u2202s \u2202x f (x) + k sign (s) ) (6.8) where k is a positive constant that verifies ss\u0307 = s (\u2212k sign (s)) = \u2212k |s| \u2264 \u2212\u03b7 |s| \u21d4 k \u2265 \u03b7 (6.9) and sign is the signum function defined by sign(s) = \u2223 \u2223 \u2223 \u2223 \u2223 \u2223 1 if s > 0 0 if s = 0 \u22121 if s < 0 (6.10) Nevertheless, the discontinuity of the signum function in the vicinity of the sliding surface s = 0 involves an infinite frequency commutation of the control law which cannot exist in practice because of switching imperfections. This leads to the so-called chattering phenomenon (Fig. 6.1) which can degrade performances of the controlled system, excite disregarded high-frequency dynamics and even deteriorate the control member. To attenuate the chatter effect, several approaches were proposed in literature. The main idea of the first approach is to substitute signum function by a continuous one such as saturation function defined by: sat(s,\u03d5) = \u2223 \u2223 \u2223 \u2223 \u2223 \u2223 \u2223 \u2223 s \u03d5 if \u2223 \u2223 \u2223 \u2223 s \u03d5 \u2223 \u2223 \u2223 \u2223 \u2264 1 sign(s) if \u2223 \u2223 \u2223 \u2223 s \u03d5 \u2223 \u2223 \u2223 \u2223 > 1 (6.11) The continuous function will ensure the convergence of system state trajectory to a thin boundary layer in the vicinity of the sliding surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001136_s00170-017-0625-2-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001136_s00170-017-0625-2-Figure5-1.png",
        "caption": "Fig. 5 Schematic of the method for recording thermal history",
        "texts": [
            " The overlap ratio for each adjacent bead is 0.5, and the beads in a particular layer are always deposited exactly over the corresponding beads of the previous layer, as shown in Fig. 4b. Herein, the \u201coverlap ratio\u201d is recorded as OVr, and calculated using Eq. 1: OVr \u00bc OL=W \u00f01\u00de where OL is overlap width and W is single bead width, as shown in Fig. 4b. Furthermore, for recording the temperature history of the molten pool in the deposition process, infrared thermography was applied. The corresponding method is illustrated in Fig. 5. The FLIRX6530sc was the infrared thermo-graph used in this study. Figure 6 shows the different sections for metallographic observation, which are represented as sections X and Y. Section X is normal to the scanning direction, while section Y is parallel to the scanning direction. All the cross-sections were made by wire cutting. The metallographic samples were mounted, ground, and polished. The sample sections were electrically etched by 10% oxalic acid and 3\u20134 V electrolysis voltage for about 30 s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002095_j.jsv.2015.02.010-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002095_j.jsv.2015.02.010-Figure7-1.png",
        "caption": "Fig. 7. Vertical forces and displacements for beam and ball.",
        "texts": [
            " In order to illustrate the effect of contact damping on the resulting beam vibrations, a parametric analysis is performed, where three values of the ad hoc viscous contact damping coefficients are considered cc \u00bc 0;20;200\u00bd N s=m. The underlying physics of the viscous damping term might be air-squeezing in the contact region giving a viscous damping effect which is acting also when there is some loss of contact. The steel ball is modelled as a rigid mass moving along the length of the beam. With reference to Fig. 7, the equation of vertical motion for the ball is therefore stated as m2 \u20ac\u03b72\u00f0t\u00de \u00bc Fc\u00f0r\u00f0t\u00de; _r\u00f0t\u00de\u00de m2g; (21) wherem2g is the weight of the ball. Beam and ball equations of motion \u2013 Eqs. (19) and (21), respectively \u2013 are combined in a set of first-order differential equations _z \u00bc Az\u00feF\u00f0z; t\u00de (22) with the initial conditions in Eq. (20). Herein, the vector z contains displacements and velocities for the ball and the beam modal coefficients. Beam and ball displacements are then found by solving the state-space system using Matlab's built-in solver ode45, which solves the system of differential equations using a variable step Runge\u2013Kutta method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003705_ecce.2018.8558256-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003705_ecce.2018.8558256-Figure5-1.png",
        "caption": "Fig. 5. First simulation scheme, with the FE field solution and the related equivalent circuit. The current isd creates the main machine flux. Torque currents along the q\u2212axis are isq = \u2212irq . The FO condition is not satisfied, \u03bbrq = 0.",
        "texts": [
            " The RFO condition and the motor performances, will be achieved using a two steps procedure, as an alternative to the iterative method. In Fig. 4 the proposed procedure scheme is shown. A first simulation is necessary to bring to zero the rotor q-axis flux, then, in the second simulation the IM FEA is closely linked with the machine RFO model, and the performances are computed. In the first simulation of the procedure, the T-form equivalent circuit is considered to compute the machine inductances (see Fig. 5). Current vectors forced in the stator and rotor windings are: is = isd + jisq; ir,1 = jirq = \u2212jisq (18) In the generic dq synchronous reference frame, stator flux linkages are analytically derived as: \u03bbsd = Lsisd +Mird \u03bbsq = Lsisq +Mirq (19) and the rotor flux linkages are: \u03bbrd = Lrird +Misd \u03bbrq = Lrirq +Misq (20) Imposing current vectors in (18), since irq = \u2212isq and ird is set to zero, machine inductances are derived as: Ls = \u03bbsd isd ; M = \u03bbrd isd ; Lls = \u03bbsd isq ; Llr = \u03bbrq irq (21) Once the field problem is solved, stator and rotor flux linkage components are computed by means of the magnetic vector potential. From (20) the mutual inductance M and the rotor leakage inductance Llr are computed. Let\u2019s remark that, when the stator and rotor windings have the same number of effective turns, the magnetizing component of stator and rotor self inductance is equal to M . The rotor inductance is computed as: Lr = M + Llr. In Fig. 5 first simulation scheme is shown: the FE field solution together with the related equivalent circuit. The rotor q-axis flux is not equal to zero, the FO hypothesis is not verified. Using magnetostatic simulations, the iron saturation is well considered. In particular, the mutual and the leakage inductances are computed taking into account the cross saturation between d and q magnetic paths and the high local saturation, that occur during on-load operation. The load-dependent inductances are computed for each set of d and q currents considered, in order to get the right value in each working point",
            " The rotor induced current can be considered uniformly distributed within the rotor bars only when the motor is working at low slip frequency, e.g. the normal operating condition, fed by the grid. In the example reported in Fig. 6, the FE field solution is shown when currents vector in (22) are imposed in the stator and rotor windings. The currents isd and isq are the same as in the the first step. In the second FEA of the procedure, the relationship (11) is applied. Looking at RFO field solution in Fig. 6, the rotor flux lines are almost parallel to the d-axis, and the difference with the field map in Fig. 5 can be observed, in which the rotor flux lines widely cross the d-axis. Thus, the rotor q-axis flux has been significantly reduced (in that example the torque current is three times the rated one). In Table I FEA results are reported in several working points, increasing the motor torque current. In each of them, using the proposed strategy, \u03bbrq,2, in the second simulation is almost equal to zero. The flux \u03bbrq,1, computed in the first step, is reported in the table to show the convergence. From the second simulation field solution stator and rotor dq flux linkages \u03bbsd, \u03bbsq , \u03bbrd and \u03bbrq are computed using the magnetic vector potential"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002666_icra.2016.7487489-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002666_icra.2016.7487489-Figure1-1.png",
        "caption": "Fig. 1: Frames of reference. N = {O0,n1,n2,n3} is the inertial frame, B = {OB, b1, b2, b3} is the body frame, r indicates the position of the flyer\u2019s center of mass from the inertial origin O0 and {\u21261,\u21262,\u21263,\u21264} are the angular speeds of the rotors.",
        "texts": [
            " In this section, we discuss the two dynamical models used to synthesize controllers for the quadrotor, which operates in two distinctive modes, multi-flip aerobatic flight and regular flight. The first model is a complex nonlinear dynamical description of the flyer. The second model assumes small attitude angles and slow angular velocities to linearize the nonlinear dynamics of the first model. Both models are used in the implementation of multi-flip aerobatic trajectories, defined as flying processes which start at a stable hover, transition to a multi-flip aerobatic maneuver, and return to a stable hover. The quadrotor employed in this research is shown in Fig. 1. To determine the dynamics of this MUAV, we define the inertial frame N = {O0,n1,n2,n3} and the body-fixed frame B = {OB, b1, b2, b3}, which has its origin and axes coinciding with the center of mass and principal axes of the MUAV, respectively. The frames N and B are related to each other through the Euler angles {\u03c8, \u03b8, \u03c6} in the Z-Y - X convention with the origin of B, OB, at position r from the inertial origin O0. Thrust is generated by four rotating propellers with angular speeds {\u21261,\u21262,\u21263,\u21264}",
            " The multiflip phase ends when the flipping angle reaches an a priori defined target angle, typically 2n\u03c0, where n is referred to as the number of flips. In most cases, the angular velocity of the quadrotor is not zero at the end of the multi-flip phase, and for this reason, the descent and re-stabilization phase is primarily used to re-stabilize the robot\u2019s attitude and stop its rapid fall. Clearly, height and time impose very stringent constraints on the flyer capabilities, which is especially true for the robot employed here (the Crazyflie shown in Fig. 1) because it has a very low thrust-to-weight ratio of 1.7. For the experiments discussed in this paper we use a Crazyflie 1.0 nano-quadrotor manufactured by Bitcraze AB [19] (shown in Fig. 1), which has a PTPT distance of 13 cm, weighs 19 g (including the battery) and a flight endurance of approximately 7 min, achieved with a 170 mAh Li-Po battery. This robot carries several on-board sensors used in the implementation of controllers, including a 3-axis accelerometer, a 3-axis gyroscope and a MEMS barometer. The MEMS barometer measures the altitude of the robot while the accelerometer and gyroscope are used to estimate the robot\u2019s attitude and generate the angular velocity signal used as feedback"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000122_921-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000122_921-Figure7-1.png",
        "caption": "Figure 7. Comparison of the position-versus-time plots of the motion of a sphere along the inclined plane and the brachistochrone.",
        "texts": [
            "5 \u2248 5a, T0 \u2248 1.1 s) and decreases for shorter ones (L = 18 \u00b1 0.5 cm \u2248 2a, T0 \u2248 0.85). In figure 6 position-versus-time plots for a sphere moving along a cycloidal track are reported. Comparison with theoretical predictions (continuous line) shows a good agreement: the graphs confirm that the time of descent does not depend on the starting height. Its value, t \u2248 0.625 s, is coherent with a cycloidal path generated by a circle of radius a \u2248 40 cm, according to the relation t = \u03c0 \u221a a g . The apparatus (figure 7, left) was used to measure the times of descent and the position versus time during the motion of a rolling steel ball for identical height drops on both the cycloid and the inclined plane. For each height the shorter time was on the cycloidal track (t = 0.30 \u00b1 0.03 s, in agreement with the theoretical value, t = 0.315 for a cycloidal track generated by a circle of radius a \u2248 10 cm). Figure 7 (right) shows the position-versus-time data for the two balls, highlighting that the time of descent is lower for the cycloid curve and the position-versus-time plot deviates from the quadratic behaviour followed by the ball moving on the inclined plane. This is in agreement with the brachistochrone property of the cycloid. Apparatus and demonstrations derived from historical experiments provide a suitable introduction to experimental and theoretical problems for physics students [28]. In this paper the experiments are related to Huygens\u2019s studies on the cycloidal pendulum and aim to explore the brachistochrone and tautochrone properties of the cycloidal trajectories"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000251_icar.2013.6766543-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000251_icar.2013.6766543-Figure6-1.png",
        "caption": "Fig. 6: Transportation using two mobile robots",
        "texts": [
            " Therefore the resultant force acting to the whole object including robots and a transported object is derived from, F = (Fk + Fck ) k=1 n \u2211 (12) An addition to the force we can calculate the resultant moment acting to the whole object as, M = lk \u00d7 Fk + lck \u00d7 Fck k=1 n \u2211 k=1 n \u2211 (13) By using these resultant forces and moment, a set of robots and an object can be treated as a single rigid object but any points we have taken into account in this section such as the robot and the midpoint between the robots, can avoid collision with the environment. Then the reference velocity vector for the whole object is derived by as follows. &x &y &z \u23a1 \u23a3 \u23a2 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 \u23a5 = \u03b1Fx \u03b1Fy \u03b2M \u23a1 \u23a3 \u23a2 \u23a2 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 \u23a5 \u23a5 (14) IV. EXPERIMENTS To verify the proposed cooperative transporting method, experiments with two omnidirectional robots and a single object are performed which schematic is sown in Fig.6. By the proposed method, potentials are defined and calculated at the center of the robots and the midpoint of the two robots. From eq.(12), the resultant force is derived as, F = (F1 + F2 )+ Fc1 (15) Also the resultant moment is described as, M = L 2 (F1x \u2212 F2 x )sin\u03c6 \u2212 L 2 (F1y \u2212 F2 y )cos\u03c6 (16) Each robot motion is created to realize whole object motions from eq.(15) and (16) with maintaining the geometric relationships between the robots and a transported object. Figure 7 shows an indoor map for the experiments. The initial position was located in the room 213. A target point was set on the mid point of a corridor. In the room, a desk existed as shown in the right bottom of the Fig.7. To avoid the local minima, the sub - target point was also set at a midpoint of a door opening. A robot formation was as shown in Fig.6, in which a distance between the robots was 1600mm then the midpoint was 800mm point The width of the door opening is about 1400mm which is not wide enough for the two robots to pass through simultaneously. To create a reference trajectory, we define the potential power on the field. By using eqs.(4), (8)-(10), potential power on the field can be derived as shown in the Fig.8. From the potential power, forces and moments to the robot system can be calculated. From these set, trajectories of robots and an object can be derived as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001075_s12206-017-0408-6-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001075_s12206-017-0408-6-Figure1-1.png",
        "caption": "Fig. 1. Cross section area of the cylinder contained two liquids.",
        "texts": [
            " In this study, based on the quasi two-dimensional flow model, the stability of spinning composite shell partially filled with two liquid phases is analyzed. According to the coupled fluid-structure model obtained by combining dynamic equations of the cylinder with the fluid, the stability status of the whirling shell is determined. It should be noted that the critical rotating frequency, mentioned in the nomenclature, is the beginning of instability (loss of dynamic stability). 2. Fluid dynamics As can be seen in Fig. 1, a whirling shell containing two ideal liquids is considered. The cylinder rotates with a fixed angular speed\u2126 . With assumption of a short shell, there is not any eccentricity. Therefore the axis is fixed and remains straight line. Rotation of shell results to the annular form of liquids. If the length to diameter of the rotor is greater than unit, Tao and Zhang [8] illustrated that the real flow inside the container can be replaced by a quasi-two-dimensional fluid. Neglecting the gravity, the linearized Navier Stokes equations in small displacement of the fluid in the rotating frame r and\u03b8 (Fig. 1) for liquid (1) can be written as follows: ( ) ( ) ( ) ( )1 1 1 2 1 1 2\u2126 \u2126r P r t r \u03b8 \u03c5 \u03c5 \u03c1 \u2202 \u2202 \u2212 \u2212 = \u2212 \u2202 \u2202 (1) ( ) ( ) ( ) ( )1 1 1 1 1 2\u2126 r P t r \u03b8\u03c5 \u03c5 \u03b8\u03c1 \u2202 \u2202 + = \u2212 \u2202 \u2202 (2) ( )( ) ( )( )1 1 0 rr r \u03b8\u03c5 \u03c5 \u03b8 \u2202 \u2202 + = \u2202 \u2202 (3) where \u03c1 shows the liquid density, and r \u03c5 and \u03b8\u03c5 refer to the radial and tangential components of fluid speed, respectively. These equations can be accounted for liquid 2, as well. The small pressure oscillations and centrifugal acceleration of the fluid make the total hydrodynamic pressure \u2018P\u2019 as follows: ( ) ( ) ( ) ( ) ( ) 1 2 2 2 1 1 \u2126 , , 2 i i m r r P p r t r r r \u03c1 \u03b8 \u2212 = + < < (4) ( ) ( ) ( ) ( ) ( ) ( ) ( ) 1 22 2 2 2 2 2 2 2 \u2126 \u2126 2 2 , , ",
            " Coupled fluid-structure system To drive the coupled relations of the liquid motions and rotating shell, the liquid boundary conditions to each other and to the cylinder wall are considered. The cylinder vibrations resulting from the external pressure on the shell is considered as loads created by liquid motions. The effect of liquid (2) on the wall of cylinder can be represent as: ( ) 0 0 2 0|r r r w w\u03c5 = + = \u027a . (45) By substituting the expansion series of and , one can obtain ( ) ( ) 0 0 2 \u02c6 \u02c6| s nr r r w n n w\u03c5 \u03b7= + = . (46) As seen in Fig. 1, the internal surface of the fluid (1) is de- fined by i r r \u03c2= + , where\u03c2 shows the small displacement of internal surface. By Eliminating the nonlinear terms and combination of equation ( ) ( )1\u02c6 n p + ( )1\u03c1 2\u2126 i r \u03c2 = 0 and ( ) ( )1\u02c6 nr \u03c5 - / t\u03c2\u2202 \u2202 = 0, the boundary condition in the internal surface for liquid (1) is derived as: . (47) Because of using ideal liquids, the boundary conditions between two liquids are based on having the same pressure and the same radial velocity as follows ( ) ( ) m m 1 2 | |r r r rP P= == (48) ( ) ( ) m m 1 2 | |r r r r r r\u03c5 \u03c5= == "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002411_ssd.2015.7348200-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002411_ssd.2015.7348200-Figure5-1.png",
        "caption": "Fig. 5. Spiral motion.",
        "texts": [
            " When the robot has reached the initial position, then the proposed algorithm will be activated and the robot will start to insert the plug-notch in the socket-key. In this phase, the perpendicular axes to the socket will be pure force controlled. The desired value of the contact force is chosen to Fzd = \u221225N . The desired values for x and y-direction are set to zero (Fxd = 0, Fyd = 0). In these two axes, force control will be additional superimposed by a spiral motion, which is illustrated in Fig. 5. This connection between force and position control may be understood as the approach of parallel position/force control introduces in ([12]). During the spiral motion, the system tests if the notch of the plug is inserted in the key of the socket. In the case that the current end-effector position in z-direction differs more than 1 cm from the z-coordinate of the first environment contact, we assumed that the notch is inserted into the key. Otherwise, if the notch is outside of the key, the diameter of the spiral motion should be increased"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002287_iet-cta.2014.0576-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002287_iet-cta.2014.0576-Figure7-1.png",
        "caption": "Fig. 7 eZ430-RF2500 WSN node used in the experimental setup",
        "texts": [
            "1479 \u22122 0.6588 1.1615 1.6823 2 \u22122.2328 \u22120.7926 \u22120.5344 \u22122 \u22120.7672 0.7926 \u22122.4656 \u22122 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 (32) IET Control Theory Appl., 2015, Vol. 9, Iss. 6, pp. 915\u2013928 923 doi: 10.1049/iet-cta.2014.0576 \u00a9 The Institution of Engineering and Technology 2015 We select eZ430-RF2500 wireless sensor nodes because they offer a fairly short time window of activity for transmitting/receiving a message [32]. In addition, these nodes are quite affordable and have modest energy requirements [33]. Each eZ430-RF2500 node (Fig. 7) is a very low-power wireless platform built around an MSP430 microcontroller, CC2500 transceiver and Digitally Controlled Oscillator serving as the internal clock. On the other hand, due to simplicity of the wireless sensor network (WSN) nodes, the topology discovery algorithm [23] is not yet implemented and time-varying topologies are not examined experimentally. For numerical simulations involving switching topologies, albeit without agent partitions, refer to [8]. To characterise the energy consumption of the eZ430RF2500, the node was connected to a laboratory power source providing 3 V and a 10 resistor was connected in series to the node"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002547_1.4033101-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002547_1.4033101-Figure2-1.png",
        "caption": "Fig. 2 Measurement of the rotor motion",
        "texts": [
            " The oil passes through the turbocharger main housing lubricating the bearings and other internal components while removing heat. The pressure is regulated through a bypass and series of valves just before the oil enters the turbocharger. A heating element is located in the reservoir of the oil circulation system. These features allow for control of both oil inlet temperature and pressure enabling operation under various conditions. Rotor motion is monitored using noncontact displacement sensors. Figure 2 depicts the sensors mounted in the compressor housing oriented perpendicular to the rotational axis spaced by a 90 deg angle. This configuration allows the rotor orbit to be captured as a function of rotational speed. The turbocharger rotor is supported by an angular contact ball bearing cartridge specifically designed for high-speed applications. In order to predict the dynamics of the turbocharger system, a coupled rotor\u2013cartridge model was developed. A description of a shaft supported by two separate deep groove ball bearings is given by Brouwer et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001502_j.ceramint.2018.07.144-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001502_j.ceramint.2018.07.144-Figure3-1.png",
        "caption": "Fig. 3. Image of powder in rotating drum during flowability measurement.",
        "texts": [
            " Table 1 shows the composition of 50 vol% alumina and 50 vol% spodumene. Due to the highly localised melting process inherent in the SLM process, alumina and spodumene powders were tumble mixed in equal amount (vol%) to ensure compositional heterogeneity between and within each deposited layer. The mixture containing alumina (spherical) and spodumene (irregular shape), which was prepared by tumbler mixing (Inversina) for 5 h is as depicted in Fig. 2. Flowability test conducted to measure the avalanche angle \u03b8 using the Revolution powder analyser is as shown in Fig. 3. The measured \u03b8 was 53.4\u00b0 (> 30\u00b0), and therefore implying poor powder flowability, can affect deposition homogeneity on the substrate plate and lead to poor mechanical properties of the parts [32]. In addition, as the spodumene powder is only larger than 100 \u00b5m in the longitudinal direction, no significant difficulties were faced during deposition at LT100 or less. Furthermore, the movement of the recoater ensures that the surface of the powder bed remains flat. Energy dispersive x-ray spectroscopy (EDS) was conducted to determine the homogeneity of the deposited layer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000621_s12206-015-0715-8-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000621_s12206-015-0715-8-Figure5-1.png",
        "caption": "Fig. 5. Experiment rig.",
        "texts": [
            " Kurtosis is widely used as a measure in machinery condition monitoring; for example, early damage in the rolling elements of a machinery often results in vibrator signals, the kurtosis value of which significantly increases because of faults in such rotating system [17]. The gears used in the experiment consist of a healthy gear and gears with three crack degrees: small crack, big crack, and tooth cut. The cracked gear conditions are shown in Fig. 4. The experiment rig consists of a 0.75 KW DC motor and a single-stage spur gear on two parallel shafts, as shown in Fig. 5. The crack is located on a driving gear with 24 teeth while the driving gear has 25 teeth. Table 1 presents the detailed experiment specifications. A speed controller connected to the input shaft controls the rotation speed. Torque load is provided to the gearbox from the eddy current magnetic brake. The device has a maximum torque capacity of 12 Nm. Vibration signals are acquired by using a Bruel and Kjaer 4506 triaxial accelerometer on the input shaft bearing housing. Signal acquisitions are conducted using an LMS Mobile and are sampled at 8"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002460_amm.805.205-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002460_amm.805.205-Figure1-1.png",
        "caption": "Figure 1: Schematic representation of the process chamber of a selective laser melting machine [11].",
        "texts": [
            " Building on a base plate, the part is grown layer-by-layer, enabling the fabrication of complex parts without geometry specific tooling. A continuous shield gas flow in the build chamber, which circulates within the machine, prevents oxidation during material melting. A broad portfolio of metals including Ti- and Al alloys, tool steel, stainless steel, super alloys as well as precious metals can be processed by the selective laser melting technology [9]. The MTT SLM 250 is equipped with a 400 W laser source and uses argon as shielding gas (see Fig. 1). The laser beam is projected onto the powder bed by an optical scanner that is separated from the build volume by a shielding glass. A cooling circuit with an external cooling unit is used to control the laser source temperature. Some of the control actuators, e. g. valves, are powered by compressed air. Before the start of the melting process, the build volume is flooded with argon until the oxygen level is reduced below 0.2 %. In order to reduce the thermal stresses within the part, the base plate is preheated before the process, lifting the temperature level closer to the material specific melting point",
            " database have been peer-reviewed. The results are presented in the sequence of the CO2PE! methodology description [6]. Goal and scope definition. The goal and scope phase consists of the steps analyzing the machine architecture, defining the system boundaries and the functional unit as well as identifying relevant process parameters. The goal of this study was to quantify the resource consumption of the SLM process by using a MTT SLM 250 machine. Its architecture was already described in detail above (see Fig. 1). The system boundaries were set closely around the machine including the external cooling unit (see Fig. 2). That implies that neither the powder generation nor the powder recycling process or post processing steps of the intermediate product were taken into consideration. In alignment with the majority of previously published studies on the selective laser melting process, the weight of the finished part measured in kilogram was chosen as the functional unit. Important parameters of the process are the build volume V and the maximum height of the part [7]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001585_aim.2018.8452392-Figure20-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001585_aim.2018.8452392-Figure20-1.png",
        "caption": "Figure 20. Peristaltic crawling robot with load cell",
        "texts": [
            " Thus, there is a problem that the traction force is insufficient and the vehicle cannot run. Therefore, by measuring the load on the robot with the load cell mounted at the rear end, feeding the data back to the microcomputer, and adjusting the opening/closing time of the solenoid valve, it is possible to suppress further decrease of the expansion diameter. This allows the robot to securely grasp the pipe wall. We aim to achieve reliable driving by demonstrating sufficient traction. The structure of the entire device is shown in Fig.19, and an external view is shown in Fig.20. In this chapter, a 4 unit type peristaltic motion robot is used. In this chapter, the time until the next operation of the solenoid valve was changed, and changes in the pulling force were measured. In the experiment, a digital force gauge was attached to the rear end of the robot and fixed to the rear end of the pipe. We then measured the traction. Changes in the pulling force were measured when the solenoid valve opening time is gradually increased from 200 ms to 800 ms in 200 ms increments. A schematic diagram of the experimental system is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000315_j.wear.2014.11.014-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000315_j.wear.2014.11.014-Figure2-1.png",
        "caption": "Fig. 2. Overview of MPR and a detail of the test zone; contact width of the roller.",
        "texts": [
            " The danger of these additives, especially EP and AW, is that if chemical activity is uncontrolled, its effects may not be localized to the asperities of the contact surface and may extensively attack the surface in high temperature and high pressure conditions. In addition to studying the effect on surface fatigue, the influence of additives on friction and wear of the contact were also analyzed. For this analysis, a commercial test machine Micro Pitting Rig (MPR) by PCS Instruments was used (www.pcs-instruments.com). A MPR disc machine, shown in Fig. 2, was used for the tests. This machine is for testing surface fatigue and also allows the measurement of the friction coefficient and wear in lubricated line contacts between discs. The MPR is currently used by the Industry to look at both macro- and micro-pitting resistance of lubricating oils on gears and bearings [14]. The equipment consists of three rings which rotate at the same velocity and in contact with a central roller connected to a separate axis, which permits rotating at different velocity from the rings. This way you can set the average (rolling) velocity um and the slide-to-roll ratio (SRR) of the contact. SRR is defined as the ratio of the sliding velocity \u0394u to the average velocity um, expressed as a percentage. The zone of interest is the rolling track of the testing roller, due to the higher strength of the outer rings [15]. The rolling track of the roller has an initial contact width of 1 mm, as can be seen in Fig. 2. This geometry of the roller produces widening of the rolling track in case of wear. The normal load W of the contact is applied on the upper ring against the roller and lower rings. The arrangement is such that the load is evenly distributed among them. The contact area is lubricated by a thermostatic bath with controllable temperature. A temperature probe is located into the test chamber with the tip of the probe close to the contact region, as shown in Fig. 2. In addition, it is equipped with a commercial ICP accelerometer (PCB, model M353B16) mounted on the test head, which is used to measure the vibration in the contact and provide real time information on the condition of the specimen's surface. The rings and roller are made of steel 16MnCr5, equivalent to F1516, which is used in the construction of gears, pinions, and cemented parts. It has undergone carburizing and quenching treatments to achieve hardness of around 780 HV in the rings and 680 HV in the roller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001592_tie.2018.2868027-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001592_tie.2018.2868027-Figure11-1.png",
        "caption": "Fig. 11. (a) Body\u2019s reciprocation during walking; (b) schematic of the BEHS [3].",
        "texts": [
            " DYNAMIC MODEL AND EQUIVALENT OF THE BEHS In this section, the BEHS is used as an example to show how the proposed equivalent circuit for the TROMAG can be applied to more sophisticated systems. The BEHS proposed in [3] can harness power from the backpack\u2019s reciprocation while reducing the metabolic cost of walking and risk of orthopedic injuries. The idea behind the BEHS is to create a phase shift between the body\u2019s reciprocation and the backpack\u2019s reciprocation through a decoupling spring: The system, shown in Fig. 11, consists of a main frame and a load 0278-0046 (c) 2018 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. frame. The main frame is firmly attached to the body, reciprocating with it during walking. The load frame, which carries the backpack, is attached to the main frame through a spring. The relative motion between the two frames can drive a generator, harvesting electricity [3]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000081_12.2084733-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000081_12.2084733-Figure2-1.png",
        "caption": "Figure 2. Fixed-free n-lamina curved beam.",
        "texts": [
            " Modeling of piezoelectric unimorphs with an initially flat shape was pioneered by Smits and Choi 1 , whose work was later modified by Weinberg 2 , who expanded the model to include multilayer actuators, and also by Dunsch and Breuguest 3 , who developed a unified approach for various loading and boundary conditions. The following modeling borrows from these analyses, using curved beam theory to account for the initial curvature, and the Maxwell strain equation to describe the strain of DE material under applied voltage. Consider a curved beam consisting of n laminae, with subtended angle  and radius of curvature  , shown in Figure 2. The beam is fixed on one end, and the free end is subjected to a moment M . Deformation of the beam is characterized by the change in subtended angle  . The general expression for strain in the \u201ci-th\u201d lamina (1) is the sum of the elastic strain, and Maxwell strain induced by an electric field (2), where  is Poisson\u2019s ratio, 0 is vacuum permittivity, r is relative dielectric constant, E is the applied electric field, and Y is Young\u2019s modulus. It is assumed that the electric field can be approximated by two parallel plates, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001204_icma.2017.8015804-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001204_icma.2017.8015804-Figure9-1.png",
        "caption": "Fig 9. Protecting container.",
        "texts": [
            " Other pins are connected to the external actuators such motors and lights. For controlling the principal motor and its speed, pulse width modulation (PWM) signal was used depending on the data received. Battery is connected to a 5v regulator for feeding the circuit, and is connected to a 12 volts motor driver, for driving the electric DC motor, as we can see in the Figure 8. A container was manufactured with the objective of protecting the circuit boards and prevent cable ruptures and malfunctions, as figure 9 shown. The circuits were handmade with a method that uses ferric per chloride, it reacts with the cooper and cleans it leaving only the surfaces covered by other material, as the toner left by a laser printer, which was used in this case, the process is shown in figures 10 and 11. With the construction of all the subsystems a procedure to test them was stablished under the regulations of the complex system, UGV specifications and UGV Design, the results are shown in Table III. F. UGV Modules Integration The integration of all the modules was relatively easy, it has no much complication due the previous validation in communication and design"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure1-1.png",
        "caption": "Fig. 1. Workspace of a redundant 3R serial robot with l 1 = 17 . 3 , l 2 = 7 . 8 and l 3 = 4 . 5 . The first joint angle is subject to joint limits \u03b81 \u2208 [15, 165] \u00b0; the second and third joints can freely rotate. This example is very similar to an example presented in [8] .",
        "texts": [
            " Typically, \u03b8 j are the lengths of linear actuators or the rotated angles of revolute actuators, whereas t i define the position and/or orientation of some output link of the robot which is of interest for a given task. The workspace can be defined as the set of values that vector t can attain subject to the kinematics of the robot (i.e., the mathematical relationship between \u03b8 and t ) and to other kinematic constraints (e.g., joint limits or avoidance of collisions). When m < n , it is said that the robot is kinematically redundant 1 [23] , which means that more DOFs than necessary are used to perform the task. Fig. 1 shows an example of redundant robot: a serial 3R planar robot in which n = 3 actuated \u2217 Corresponding author. E-mail addresses: apeidro@umh.es (A. Peidr\u00f3), lpaya@umh.es (L. Pay\u00e1). 1 In the following, we will omit the word \u201ckinematically\u201d, assuming that all the redundancies mentioned in this paper are of this type. https://doi.org/10.1016/j.mechmachtheory.2018.05.013 0094-114X/\u00a9 2018 Published by Elsevier Ltd. revolute joints control the Cartesian coordinates t = [ t 1 , t 2 ] = [ x, y ] of its tip in the plane ( m = 2 )",
            " For example, joint limits (which are usually modeled as inequalities) can be easily rewritten as equalities [19] [even as quadratic equations [8] ] by introducing auxiliary variables. But this is difficult in general for more complex constraints, like the prohibition of mechanical interferences (collisions). Usually, most of the methods previously described only obtain the boundaries of the workspace, omitting highly valuable information of its internal structure. It is well known that inside the boundaries of the workspace there may exist interior barriers (see Fig. 1 ) which imply motion impediments for the robot [9] . Knowing the distribution of such barriers inside the workspace is necessary for effectively planning trajectories in the task space, since a given trajectory that crosses one of these barriers may be unfeasible, depending on the values of the joint coordinates when approaching the barrier (see Section 2 ). Singularity-based methods [8] naturally obtain the interior barriers among the solutions of the aforementioned system S . Geometrical methods may also be able to identify such interior barriers in simple cases [see the comments on [32] in Section 2",
            " However, to the best of our knowledge, these methods have not been used for obtaining interior barriers in redundant robots under complex collision constraints. Considering the advantages and limitations of all previous methods, sampling methods are promising for obtaining the interior barriers of the workspace under complex collision constraints, due to their ability to easily handle these constraints. FK-based sampling methods might reveal the interior barriers in simple cases if joint coordinates are sampled from appropriate random distributions [37] . For example: in Fig. 1 , one can check that, if \u03b81 is sampled from a U-shaped beta distribution in [15 \u00b0, 165 \u00b0] [11] and { \u03b82 , \u03b83 } are uniformly sampled in [0, 360 \u00b0], then the density of randomly generated task points will be higher near the interior barriers, revealing them. However, this method is not completely robust nor predictable, and not easy to generalize. Therefore, the last available option is to detect interior barriers using IK-based sampling methods. In non-redundant robots, the solutions of the inverse kinematics for a given task-space point generically are a finite number of isolated points of the joint space",
            " For each query point t q , its preimage self-motion manifolds are approximated by the set M (t q ) of all sampled joint coordinate points \u03b8i that have generated task positions t i near t q when solving the forward kinematics during the previous randomly-sampling stage. Then, all joint samples contained in M (t q ) are clustered by solving a Minimum Spanning Tree problem [14] , which allows for the identification of disjoint self-motion manifolds. Demers [16] used this method to identify the Jacobian surfaces that separate neighboring w-sheets in the workspace of serial 3R redundant robots, in the absence of kinematic constraints (although the method may easily accommodate such constraints since it is a sampling method). To this end, radial coordinate \u03c1 ( Fig. 1 ) was swept and self-motion manifolds were estimated at some discrete values of \u03c1 , following the procedure described above. Then, Jacobian surfaces were identified with the values of \u03c1 at which the number of disjoint self-motion manifolds identified by the clustering method changed. In the absence of kinematic constraints, the changes in the number of disjoint self-motion manifolds can be used to reliably identify Jacobian surfaces [10] . However, the focus in the present work is on barriers, and the changes in the number of manifolds when moving along the task space do not necessarily imply the vanishing of manifolds (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003296_physrevfluids.2.123903-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003296_physrevfluids.2.123903-Figure2-1.png",
        "caption": "FIG. 2. (a) Photograph of the experimental apparatus used to measure the drag force on a grooved cylindrical sample under aerodynamic loading. The sample (1), mounted across the wind tunnel (2), is rigidly connected to a force sensor (5) via an air bearing (4) to measure the exerted aerodynamic drag forces. The internal pressure of the sample is set by a regulated vacuum pressure source (3). The air flow with average far-field speed U is aligned perpendicularly to the axis of the sample. (b) Schematic diagram of the cross section of the sample (top) comprising a rigid acrylic skeleton (red) covered by a thin latex film (blue, thickness t = 0.25 mm). Depressurizing ( p) the inner cavity of the shell results in deeper grooves (bottom).",
        "texts": [
            " These samples were then loaded aerodynamically in a flow field generated by a wind tunnel. The 123903-3 resulting drag force was measured directly by a precision system containing a load cell. The velocity of the incoming flow was then varied systematically to determine the drag coefficient, Cd, over a range of Reynolds number (2.5 < Re [\u00d7104] < 15), and for different set values of the groove depth. Next, we describe in more detail first the experimental apparatus and then the sample fabrication procedure. Figure 2(a) shows a photographs of the experimental apparatus used throughout this investigation. The incoming air flow was produced by an open return wind tunnel with a 30.5 \u00d7 30.5 cm2 test section, which was capable of producing uniform steady flow speeds of 5 < U [ms\u22121] < 34, measured by a Pitot tube and a high-accuracy capacitance manometer (690A Baratron, MKS Instruments). For all the values of U used in this study, the turbulence intensity, defined as fluctuations of the wind speed, was below 1% of the mean velocity. The flow direction was aligned with the y axis [see the definition of the axes in Fig. 2(a)]. The cylindrical samples were mounted such that they spanned the width of the wind tunnel, along the x direction, perpendicular to the incoming flow, and positioned at the vertical center of the test section. The sample protruded through holes in the y\u2212z side walls of the test section. The cylindrical samples were 43 cm long with a 3.5 cm radius (measured from the center of the cylinder to the extremity of one of the spokes). Circular caps made out of acrylic were inserted at both ends of the sample to ensure sealing. One of these end caps [left and side of Fig. 2(a)] contained a port to connect the sample, via PVC tubing, to a vacuum pump (DOA-P704-AA, Gast). A high-resolution electronic pressure control valve (QPV1, Proportion-Air, Inc.) was introduced between the sample and the vacuum pump to automatically regulate the pressure of the system using a data-acquisition device (DAQ, USB-6008, National Instruments). This pressure control valve was then controlled by a custom LABVIEW program (LABVIEW 2010, National Instruments). The two end caps of the sample were mounted onto a U-shaped aluminum frame, which was itself bolted to one end of a linear air bearing (RAB2, Nelson Air Corp",
            " With this setup, 123903-4 aerodynamic drag forces exerted onto the grooved cylindrical samples could be measured in the range 0.05 < Fd [N ] 22.2. Both the force and the wind velocity signals were digitized simultaneously by the DAQ system. The experimental setup detailed above allowed for the control of the internal pressure of the cylindrical samples (and thus the shape of the grooves), while simultaneously enabling measurement and recording of both the drag forces on the sample and the velocity of the incoming flow. In Fig. 2(b), we present a schematic diagram of the cross section of our cylindrical samples, which were custom fabricated. The samples consisted of a latex film [represented by the blue line in Fig. 2(b)] stretched over a rigid acrylic skeleton [represented by the red circle with spokes in Fig. 2(b)]. To manufacture the rigid skeleton, first, spokes with a height of 9.5 mm were laser cut (Laser Pro, Spirit GLS) out of 1.6-mm-thick acrylic plates. Second, a series of rectangular holes were laser cut on the surface of an acrylic tube (50.8 mm outer diameter and 3.175 mm wall thickness). There were two sets of holes cut into the base tube. The first set was uniformly spaced around the circumference of the cylinder and was subsequently plugged with the spokes to create the skeletal structure, with an outer diameter of 69",
            " The two elastic coefficients can be used to estimate the shear modulus, G, of the material. From the fitting of the Gent model to the stress-stretch data we obtained Jm = 37.5 \u00b1 1.7, C1 = 115.7 \u00b1 5.7 kPa, C2 = 518.0 \u00b1 30.2 kPa, and G = 577 \u00b1 32 kPa. In this section, we present the results of both mechanical experiments and finite element (FE) simulations used to characterize how the shape of the grooves; specifically, the groove depth, d, depends on the internal pressure, p. This relationship between d and p will be required later (Sec. IV) to inform the wind tunnel experiments. In Fig. 2(b), we show a schematic diagram of a segment of the latex film suspended above a single cavity of the rigid skeleton. When the value of the pressure inside the cavity, pi , is smaller than the exterior pressure, pe, the pressure differential, p = pe \u2212 pi , loads the latex film. This loading causes the film to deform inward and, thus, deepen 123903-5 the surface groove. To study the mechanics of deformation of the film, we performed a series of experiments where p was fixed to a set value and a laser sheet was then projected onto its surface, at a 45\u25e6 angle relative to the central axis of the cylinder",
            " 3(b), we show examples of the experimental and computed surface profiles of a single groove, at five different values of the internal pressure ( p = {2, 4, 6, 8, 10} kPa), for a sample with N = 14 grooves. The groove profiles closely resemble catenaries, especially for small values of p, which is to be expected given the nearly radial pneumatic loading on the latex film. We find excellent agreement between the experiments and the FEM simulations, noting that the latter involve no free fitting parameters. From the groove profiles, a representative set of which was shown in Fig. 3, we measure the groove depth as d = ro \u2212 rf , where ro and rf are defined in Fig. 2(b), and d is drawn schematically in Fig. 3(a). It is worth noting that the groove depth, d, is defined relative to the initial state of the sample with no pressure being applied. This means that for a sample with N grooves, d = 0, corresponds to an N -sided polygon rather than a perfect cylinder. However, for increasing values of N , the sample approaches the smooth cylinder case. In Fig. 4(a), we plot d versus p, for four different samples with N = {14, 16, 20, 24}, from both the experiments and the corresponding FEM simulations; excellent agreement is found between the two"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002012_msec2014-4029-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002012_msec2014-4029-Figure2-1.png",
        "caption": "FIG 2: GEOMETRY OF THE MODEL AND MESH DESIGN",
        "texts": [
            ", starting temperature from which coefficient of thermal expansion is obtained) and \ud447 is initial temperature. For simulation half symmetric model with dimension 13mm \u00d7 13mm\u00d7 80mm have been developed. Coupled analysis for thermal and mechanical phenomena is considered. To evaluate the effect of thermal stresses produced during the process 8 nodded coupled temperature-displacements, C3D8T elements are used. Uniform mesh size of 1mm\u00d71mm has been considered across the cross-section of the clad and substrate. Figure 2 presents a pictorial representation of the model geometry along with the dimensions and meshed geometry. Total number of node is 12322 and elements are 10440. In powder injection technique powder, material or mass is added continuously on to the substrate, and its modelling requires the geometry of the part to be represented by a mesh of finite elements that changes over time so as to simulate the powdered nature of the material. This is achieved by means of successive discrete activation of new set of elements into the computational domain or geometry using element birth and death feature as illustrated by Figure 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001289_roman.2017.8172295-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001289_roman.2017.8172295-Figure2-1.png",
        "caption": "Fig. 2: Swivel angle (\u03b1) representing the elbow position (e) for fixed shoulder (s) and hand (h) positions.",
        "texts": [
            " Since the forward kinematics evaluates the desired pose at a frequency of 100 Hz, the inverse kinematics output was computed at the same frequency. Each algorithm loop was computed N times before providing output, assuming N \u2217 tIK \u2264 1 100 , with tIK representing the algorithm computational time. This approach was adopted for all the three algorithms. The AJ algorithm was implemented using the swivel angle as elbow position measure [14]. This angle \u03b1 represents the elbow position for fixed shoulder and hand positions (see Fig. 2). The Augmented Jacobian JA is computed adding to the Jacobian a row which represents the relation between the time derivative of \u03b1 and the joints configuration. Also the cartesian pose (and the related v and e values) is augmented with the additional swivel angle value. q(tk+1) = q(tk)+(J\u22121 A q(tk)\u2217 (v+Ke))\u2217 (tk+1\u2212 tk) (13) JDLS A = JT A (JAJT A + k2In) \u22121 (14) The algorithm was implemented according to Eq. 13 and 14, where In is the nxn identity matrix and J\u22121 A is replaced by JDLS A in order to avoid singularity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001907_tmag.2015.2448519-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001907_tmag.2015.2448519-Figure1-1.png",
        "caption": "Fig. 1. Problem under investigation. A permanent magnet moves with a sinusoidal velocity profile over a conductive slab.",
        "texts": [
            " In these studies, they considered the interaction between the primary magnetic field of the magnet and the secondary field from the induced eddy currents, leading to field suppression and skin effect in case of high velocities. Because of its ease and inexpensive realization, it is a great illustration of Faraday\u2019s law for undergraduate students. This paper addresses the analytical calculation of motion-induced eddy currents in the case of harmonic motion of current carrying coils or permanent magnets in the vicinity of electrical conductors. The general principle is shown in Fig. 1. Despite the academic nature of this problem, analogies can be found in many modern engineering problems where oscillations occur and motion-induced eddy currents together with the associated forces are utilized. A typical example is eddy current brakes. Strong magnetic fields are used to generate drag forces that act on moving actuators [4], [5]. In contrast, in magnetic levitation (MAGLEV), one is interested to maximize the lift-to-drag ratio for increased efficiency. These kinds of transportation systems can avoid ground friction and are able to provide high-speed transportation [6]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001325_0142331217740947-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001325_0142331217740947-Figure1-1.png",
        "caption": "Figure 1. The convertible quad tilt wing (QTW) vehicle in vertical flight configuration.",
        "texts": [
            " To validate the proposed PSO-fuzzy GS-PID design approach, some comparisons are also presented. Hardware cosimulation of the designed and optimized control structures are presented, based on the proposed CAD platform, developed around an embedded NI myRIO-1900 board and LabVIEW software. Finally, conclusions are drawn. Modelling of the quad tilt wing UAV System description and referential frames The model of the studied QTW is supposed as a rigid and symmetrical six-degrees-of-freedom (6DOF) body as shown in Figure 1. With the flexibility of the wings and fuselage neglected, the aerodynamic centre and centre of gravity coincide. To establish a mathematical model of the QWT, both coordinate systems, the earth reference frame F i = Oi, xi, yi, zif g and body-fixed frame Fb = Ob, xb, yb, zbf g, are considered. The position and the attitude angles of the QTW aircraft in the earth frame are given by vectors j = x, y, z\u00f0 \u00deT 2 R 3 and h= f, u,c\u00f0 \u00deT 2 R 3, respectively. We denote by f 2 p=2,+p=2\u00bd \u00bd , u 2 p=2,+p=2\u00bd \u00bd and c 2 p,+p\u00bd \u00bd the roll, pitch and yaw Euler angles, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001621_rpj-04-2017-0057-Figure18-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001621_rpj-04-2017-0057-Figure18-1.png",
        "caption": "Figure 18 Example part basic dimensions (inches)",
        "texts": [
            " Recall, the colored D ow nl oa de d by T he U ni ve rs ity o f T ex as a t E l P as o A t 0 4: 24 1 6 O ct ob er 2 01 8 (P T ) feature mesh model is used extensively for analysis and process planning; however, all toolpath generation is done on the native CAD file in the CAM package. Hence, the accuracy of the machining process is not inherently different than conventional NC programming; at least not due tomodel input. A dimensional inspection was conducted for the machined example part using a ZEISS CalypsoTM CMM. The as-designed dimensions are given in Figure 18 and inspection results for a selected set of features are provided in Table I. The results show reasonable accuracy for a machining process, with maximumdeviation on the order of 0.0038 inch (0.09mm). The proposed method takes the advantages of an AM process to create a near net shapemodel, greatly reducing the volume of material removed, as compared to machining alone. In the previous example part, the material removal volume is reduced from 8 inch3 (from round stock) or 4 inch3 (from square stock), to less than 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003644_tr.2018.2870276-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003644_tr.2018.2870276-Figure1-1.png",
        "caption": "Fig. 1. Effects of wear on the profile of high-speed train wheels.",
        "texts": [
            " Section III provides a dynamic reprofiling policy with periodic inspections and formulates a maintenance model in the framework of the SMDP, which mainly includes state space and action space definition, transition probability calculation, model computation, and performance evaluation. Section IV provides a numerical example and analyzes the results. Section V presents the concluding remarks. As the increase of mileage for a high-speed train, the tread and flange of its wheels wear to different degrees [28]. The wear of the flange as well as the tread is the most important cause of high-speed train wheels failure. In Fig. 1, the initial profile of the wheel is represented by the solid line, while the worn profile of the wheel is represented by the dotted line. The diameters of the initial wheel and the worn wheel are represented by D and D\u2032, respectively. The initial flange and the worn flange thicknesses are represented by Sd and Sd\u2032, respectively. The values of D\u2032 and Sd\u2032 could be observed via regular inspections using on-board or wayside inspection systems. The states of the wheel are directly reflected by the detected values"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003777_s40799-019-00308-0-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003777_s40799-019-00308-0-Figure2-1.png",
        "caption": "Fig. 2 Several structural forms of slewing bearing: (a) single-row ball; (b) double-row ball; (c) cross-roller; (d) three-row roller",
        "texts": [
            " Degradation performance description . Health stage division Slewing bearings are key components extensively assembled in wind turbine generators, tunnel-boring machines, tower cranes, military technology and so on. Figure 1 depicts its various applications vividly. For diverse working conditions, there exists several structure forms of slewing bearings including but not limited to (a) single-row ball bearing; (b) double-row ball bearing; (c) cross-roller bearing; (d) three-row roller bearing which are listed in Fig. 2. In addition, it works as a large-size rolling rotational connection between slewing systems with rotational speed ranging from 0.1 rpm to 5 rpm and often bears axial force, radial force and overturning moment under harsh outdoor environments. These features are quite different from small bearings. Accordingly, traditional data-driven [1] based condition monitoring or life prediction approaches consisting of signal preprocessing and establishment of life models may not be very applicable to slewing bearings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000852_06630.0035ecst-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000852_06630.0035ecst-Figure2-1.png",
        "caption": "Figure 2. (a) The Kretschmann geometry for SEIRA. (Note that diagram is not to scale.) (b) Examples of geometries of dipole moments of the species in solution. Dipole moments parallel to the metal surface result in destructive interference while dipole moments perpendicular to the metal surface result in constructive interference.",
        "texts": [
            " In this study, we utilized SEIRA to observe dynamic structure between the IL, 1-ethyl-3-methylimidazolium bis(trifluoromethane)sulfonimide [C2MIm][TFSI] (note that the anion is also commonly called \u2018bistriflimide\u2019), and gold electrodes during cyclic voltammetric perturbation. The surface enhanced absorption effect has been described by other research groups.8,9 The reason we have chosen [C2MIm][TFSI] is because of a recent study by K. Motobayashi and co-workers wherein hysteresis at 2 mV/s scan rate was reported.10 For our purposes, this study allowed us to verify our own experimental methodology and also to further explore and interrogate the dynamic behavior of this system. Figure 2a shows a spectroelectrochemical cell in the standard Kretschmann geometry for performing SEIRA measurements.11,12 The diagram includes standard three electrode system: working electrode (WE), counter electrode (CE), and reference electrode (RE) to control and measure electrochemical events concurrent with spectroscopic assessments of interfacial dynamics. IR light is directed through an attenuated total reflectance (ATR) 36 ) unless CC License in place (see abstract).\u00a0 ecsdl.org/site/terms_use address",
            " Notable differences between our cell and that diagramed is that our system utilizes a quiescent solution (no flow system as depicted) and a two electrode setup since we only monitored relatively small non-faradaic (capacitive) current. As mentioned, the SEIRA technique takes advantage of surface plasmon-related amplification of surface-specific interactions between solution species and a conductive electrode. The enhancement phenomenon has associated surface selection rules for the vibrational absorbance observed, Figure 2b. Briefly, conduction band electrons in a relatively thin (<100 nm) nanostructured electrode strongly couple with the frequencies of impinging infrared light and confine IR energy as surface plasmons. Conduction band electrons in the electrode are also free to respond to the polar (in the case of IR-active modes) or polarizable (in the case of Raman-active modes) vibrations of adsorbed species. The resulting image charges interfere with their molecular counterparts either constructively or destructively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001682_iros.2018.8593935-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001682_iros.2018.8593935-Figure3-1.png",
        "caption": "Fig. 3: Semantic environment model of a fridge.",
        "texts": [
            " The activity parsing and recording module takes the stream of time-stamped world states together with the abstract world states, force-dynamic events and motion events from the agent and generates a symbolic activity representation stated in a first-order time interval logic. The symbolic activity representation is time synchronized with subsymbolic stream data that includes the agents and objects poses and shapes, and images. The semantic environment extraction module maps over the data structures of the world state and asserts for each relevant object and its parts symbolic names, the label category, the part-of hierarchy, the articulation chains and models, and other relevant symbolic relationships (Figure 3). The virtual image capturing module can place cameras in the game environment, access their scene\u2019s built in (deferred) rendering information, such as color, depth, specular etc. image data, and further extend it by segmenting the image into objects and labeling them with their corresponding symbolic names. Around the knowledge representation module layer is the cognitive capabilities layer. This layer includes the KnowRob query answering service, the robot perception component, a component for the mental simulation of actions, a component for learning from virtual reality demonstrations, and another for learning action models from virtual experience data",
            " Nowadays, game engines can simulate and render complex scenes with update rates of up to 90hz (typically required by VR applications), these simulated environments can be extended to large open worlds (hundreds of square kilometers) while maintaining the accuracy of being able to show realistic leaves on trees and single blades of grass on fields [9]. Finally, the computational resource requirements of physics engines have dropped such that they can run on devices such as smart phones and within modern web browsers. The combination of these developments have brought us to a point where it is easily possible to run physics engines at execution time as components of robot control programs. This makes it possible for robots to maintain a photorealistic model of their environment (see Figure 3) with approximate physics simulation. Having access to the data structures of such a model, it allows robots to retrieve detailed subsymbolic information about their world, to mentally look at scenes, and to simulate action executions. The size of the environments, the number of objects and the level of detail are modeled in a way that goes far beyond what symbolic knowledge services could provide so far. An advantage that we gain by combining game enginebased knowledge processing with symbolic knowledge processing is that robots can construct problem specific abstractions on the fly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002532_tasc.2016.2526025-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002532_tasc.2016.2526025-Figure4-1.png",
        "caption": "Fig. 4. Position of the SC coil.",
        "texts": [
            " Same as the circuit equation case, Rung\u2013Kutta method is used to solve the motion equations. The SC coil quenching occurs when temperature of the SC coil rises and the SC coil becomes resistance. As a result, the SC coil is not able to keep the superconducting state. SC coil quenching has not occurred in the experimental operation until now. But it is necessary to evaluate this case. The SC coil current iq satisfies (6) [8]: iq = i0 ( 1.0 + t\u2212 tq 0.501 ) exp ( \u2212 t\u2212 tq 0.501 ) (6) where i0 is the initial SC coil current, tq is the coil quenching time. The bogie runs at v = 120 m/s. Fig. 4 shows position of the SC coil. As the bogie is symmetrical structure, one side (SC1SC4) of the bogie is studied. Weight of the bogie balances with the levitation force (y, z) = (0,\u22120.0378) m [7]. Initial rotational position is at (\u03b8p, \u03b8yw, \u03b8r) = (\u22120.000804, 0, 0) rad. \u03b8p, \u03b8yw, and \u03b8r shows pitching, yawing and rolling angle. As no SC coil exists in front of the SC1, vertical position of the SC1 goes down a little to balance the pitching torque. Thus \u03b8p is not zero even in the steady state. Fig. 5 show running characteristics of the bogie"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000875_tmag.2015.2443133-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000875_tmag.2015.2443133-Figure1-1.png",
        "caption": "Fig. 1. Measurement device for magnetostriction. (a) Vertical single yoke. (b) Laser vibrometers locations from the top view.",
        "texts": [
            " Corresponding author: Y. Zhang (e-mail: zhangyanli_sy@ hotmail.com). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TMAG.2015.2443133 describe the vector relation between magnetic flux density vector B and principal strain vector \u03bb and is applied to the finite-element (FE) analysis. The validation of the model is verified. A device with a vertical single yoke and a laser vibrometer has been built, as shown in Fig. 1. The sample size for this device is 500 mm long (along the excitation direction) and 100 mm wide, and one end of the sample is fixed and the other is free. An optical sensor with a 10 nm/m resolution is located at the free end of the sample. In order to fulfill the measurement of the in-plane magnetostriction strain at an arbitrary direction, the optical sensor is improved and can be controlled to move to three different locations in sequence at the angles of 0\u00b0, 45\u00b0, and 90\u00b0 with respect to the excitation direction, as shown in Fig. 1(b). In addition, to investigate the anisotropic property of magnetostriction, the samples are cut at 15\u00b0 intervals from the rolling direction (RD) to the transverse direction (TD). The in-plane magnetostriction \u03bb(\u03c4, \u03b2), in an arbitrary direction, can be expressed as \u03bb(\u03c4, \u03b2) = L L = \u03b5x cos2 \u03b2 + \u03b5y sin2 \u03b2 + \u03b3xy sin \u03b2 cos \u03b2 (1) 0018-9464 \u00a9 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index",
            " By considering the complication of magnetoelastic coupling and computation cost, a simplified method of the coupling with FE analysis is developed in this paper. We carry out, first, the solution of the magnetic field and calculate the vector B in each mesh element. Then, according to (5), the \u03bbpp and \u03b8\u03bb in each element can be interpolated from the curves in Fig. 4. Finally, the magnetostriction distribution of the whole simulation model and its deformation can be obtained. Taking the measurement device with the so-called vertical single yoke in Fig. 1 as an example, its 3-D mesh is shown in Fig. 8. Using the modeling method mentioned above, the simulation results are compared with the measured ones at four different values of B = 0.612, 0.806, 1.243, and 1.419 T, shown in Fig. 9. The computed magnetostriction refers to the averaged value of magnetostriction in all elements. It can be seen that the modeled magnetostriction agrees well with the measured values. The vector magnetostrictive property between principal strain vector and magnetic flux density vector has been measured using an improved version of laser vibrometers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002969_s00170-017-0213-5-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002969_s00170-017-0213-5-Figure5-1.png",
        "caption": "Fig. 5 The Mesh geometry system",
        "texts": [],
        "surrounding_texts": [
            "Smoke visualization was applied to confirm the flow patterns and flow structures inside the working chamber using the Mie scattering technique with a strong and stable lighting source. The diode-pumped solid-state (DPSS) laser (MGL-F-532-2W, Unice E-O Services, Inc., Taoyuan, Taiwan) with a cylindrical Table 1 Computational parameters for simulating the working chamber Problem type Flow turbulence Turbulence model Standard k-epsilon Wall function Standard wall Fluid N2 Density (kg/m3) 1.1614 Viscosity (kg/m\u00b7s) 1.846 \u00d7 10\u22125 Pressure (Pa) 101,300 Inlet boundary type Velocity inlet Velocity inlet (m/s) 7 Inlet turbulence intensity for kinetic energy 0.045 Hydraulic diameter for dissipation rate (m) 0.05 Outlet boundary type Pressure outlet Grids 2,625,713 Gravity in Y-direction (m/s2) \u22129.81 Initial X-direction velocity (m/s) 1E\u2212006 Initial Y-direction velocity (m/s) 1E\u2212006 Initial Z-direction velocity (m/s) 1E\u2212006 Linear solver AMG Residual error 1 \u00d7 10\u22124 Max. iterations 2000 Min. residual 1E\u2212018 Spatial differencing method Velocity 2nd order Turbulence 2nd order Solvers sweeps 500 Inertial relaxation 0.2 Linear relaxation 0.1 Turbulence viscosity ratio 65 lens was used to create the laser sheet for different sectional flow fields. Smoke was generated by the smoke generator (Z-800IIR, Antari Fog Machine, Taoyuan, Taiwan) fed with mineral oil and operated in 0.7 kW power consumption. The flow behavior can be easily observed when the smoke passed through the tube and the trapezoid push nozzle. The images of flow visualization were recorded by IDT Vision N4 series high-speed cameras, and the resolution of the all-axes charge-coupled device (CCD) can be up to 2336 \u00d7 1728. It could record the maximum 5000 frames per second (fps). Velocity Inlet Pressure Outlet Fig. 6 The boundary conditions of the meshes"
        ]
    },
    {
        "image_filename": "designv11_13_0003126_b978-0-12-803581-8.10316-9-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003126_b978-0-12-803581-8.10316-9-Figure8-1.png",
        "caption": "Fig. 8 Light addressable potentiometric sensor (LAPS) as biosensor. Pic: pubs.rsc.org.",
        "texts": [
            " The EIS structure is scanned by an external DC bias to form accumulation, depletion and inversion layer at the interface of the insulator (SiO2) and semiconductor (Si), sequentially. Light induced charge carriers are produced when light is subjected to the silicon layer. These charge carriers are separated by the internal electric field and thus photocurrent can be detected from the circuit. A layer of Si3N4 is fabricated on the surface of LAPS as the H\u00fe -sensitive layer. When the bias voltage is constant in the middle region, change of the photocurrent indicates the pH change of the electrolyte. Another example of LAPS as a biosensor is shown in Fig. 8. A plug-based microfluidic device [16\u201318] based on LAPS can be used for measurement of continuously flowing samples. It is also possible to visualize the spatial distribution of ions inside the flow [18]. In the case of continuously flowing samples, the consumption of the solution was still large and this minimized the main advantage of microfluidic devices. To solve this problem, a plug-based version of a microfluidic device combined with LAPS is developed. Such arrangement can control the position of a plug in the channel, it can mix two plugs for reaction and it can detect the change by differential measurement [18]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000785_j.jsv.2015.06.048-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000785_j.jsv.2015.06.048-Figure3-1.png",
        "caption": "Fig. 3. Triggering phase error.",
        "texts": [
            " In this way, any changes of the parameter \u0398 (calculated with respect to the reference angular frequency \u03c9r) will be due to an angular frequency variation and not to fictitious phase angle variations. On the other hand, a triggered acquisition does not allow to trace actual phase variation of the signal. It is worth noting that the acquisition is a discrete process; even if the acquisition frequency can be really high, it is not possible to catch exactly the signal maximumwith the triggering criterion. Fig. 3 shows an example of a sample triggered with a delay of \u0394 with respect to the maximum of the signal. This would lead again to a fictitious phase angle which would be different in the sequent samples, causing an error in the frequency shift detection. The range of this error depends on the acquisition rate. The minimum value of triggering delay is \u0394min \u00bc 1 2\u03c90=F (\u03c90 being the signal starting angular frequency and F the acquisition rate), while the maximum delay is \u0394max \u00bc 3 2\u03c90=F . The parameter \u0398 will be oscillating around the mean value \u0394 \u00bc\u03c90=F with an amplitude of \u0394a \u00bc 1 2\u03c90=F "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002584_icrom.2014.6990952-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002584_icrom.2014.6990952-Figure1-1.png",
        "caption": "Fig. 1. Transporting an object by multiple robots [3].",
        "texts": [
            " INTRODUCTION ulti Robot Systems (MRSs) have been developed to carry out sophisticated tasks which a single robot is incapable of it, especially in the presence of uncertain- ty, incomplete information, and distributed control [1]. MRSs provide the possibility of doing tasks faster, cheaper, and more reliably due to their high redundancy, capacity, and flexibility [2]. Such properties caused their wide application in performing different tasks such as object transportation in various environments like manufacturing plants, warehouses, and construction sites. Fig. 1 shows transporting a heavy flat object by four mobile robots. There are two main approaches in multirobot systems: Centralized and Decentralized (or decoupled). In the first approach a central processing unit plans and controls beha- viors of the agents (robots) through a unified architecture, while in the second approach the robots are in charge of planning and coordinating individually and locally. Although the Centralized approach is complete and can provide global optimal solutions, it is vulnerable to faults and is challenged by the high dimensions of the composite configuration space of the system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001935_978-3-319-59677-8_3-Figure3.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001935_978-3-319-59677-8_3-Figure3.1-1.png",
        "caption": "Fig. 3.1 Forces and moments acting upon a three dimensional solid",
        "texts": [
            "11), we obtain sin 2\u03b5 z12 \u00bc cos \u03c7 * \u00fe2 k12 * z 1\u00fe k11 * z 1\u00fe k22 * z : Since k12 * z \u00e612z, we have sin 2\u03b5 z12 \u00bc cos \u03c7 * \u00fe2\u00e612z 1\u00fe k11 * z 1\u00fe k22 * z : Because for thin shells 1 + kiiz 1, 1\u00fe kii * z 1, the final formulas for tangent and shear deformations of Sz, take the form \u03b5 zii \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe 2\u03b5ii p 1\u00fe z kii * 1, sin 2\u03b5 z12 \u00bc 1 2 cos \u03c7 * \u00fe2\u00e612z : \u00f03:14\u00de 3.2 Forces and Moments Consider a differential element in the deformed shell bounded by the surfaces \u03b1i\u00bc const, \u03b1i+ d\u03b1i\u00bc const, and z * 0:5h (Fig. 3.1). Internal forces acting upon the element are given by p1 H2 * d\u03b12dz and p2 H1 * d\u03b11dz. Here pi are stress vectors, H2 * d\u03b12dz, H1 * d\u03b11dz are the surface areas of differential boundary elements at z \u00bc const. Integrating the internal forces over the thickness of the shell, we obtain the resultant of force vectors, R1, R2 in the form R1 \u00bc \u00f0z2 z1 p1 H2 * d\u03b12dz, R2 \u00bc \u00f0z2 z1 p2 H1 * d\u03b11dz, z1 \u00bc h=2; z2 \u00bc \u00feh=2\u00f0 \u00de: Dividing Ri by the length of linear elements, Ai * d\u03b1i \u00bc Ai ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe 2\u03b5ii p d\u03b1i (i \u00bc 1, 2), we find R1 \u00bc \u00f0z2 z1 p1 H2 * dz A2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe 2\u03b522 p , R2 \u00bc \u00f0z2 z1 p2 H1 * dz A1 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe 2\u03b511 p : \u00f03:15\u00de Similar reasoning leads to the definition of the resultant internal moment vectors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000500_s11665-012-0227-y-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000500_s11665-012-0227-y-Figure7-1.png",
        "caption": "Fig. 7 Constrains on springback model",
        "texts": [
            " Since springback is assumed to be a pure elastic unloading process, and no reverse yielding occurs, the constitutive Eq 1 for a pure elastic unloading process is carried out by user-subroutine UMAT under the platform of ABAQUS/Standard, as shown in Fig. 6. 12\u2014Volume 22(1) January 2013 Journal of Materials Engineering and Performance 5.2.2 The Boundary Condition. The boundary condition should be imposed to constrain the whole translation and the whole rotary movement of the bent-tube, to keep the system energy unchangeable during springback computation. The boundary condition should not restrict the real springback deformation at the same time. An encastre boundary is imposed at section A (as shown in Fig. 7), for there is rarely deformation at section A. According to the practical constraints of the dies, the symmetry boundary is added too. The predefined stress-strain field is also applied to the tube at the same time. Then, the springback prediction model is established. The established springback prediction model is validated experimentally. The experiments are carried out using a W27YPC-63NC microcomputer controlled bender, as shown in Fig. 8(a), and various rotary-draw bending dies are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001759_b978-0-12-800351-0.00011-0-Figure11.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001759_b978-0-12-800351-0.00011-0-Figure11.1-1.png",
        "caption": "FIGURE 11.1 Laser Treatment of Materials",
        "texts": [
            " When a biological object absorbs laser energy, the resulting nonuniform temperature distribution causes internal forces leading to thermomechanical transients and thermoelastic deformation. These changes are the driving force of all laser ablation processes that are not photochemically mediated. In the absence of photomechanical or phase transition processes, the power absorbed by the tissue in response to pulsed laser radiation is entirely converted to a temperature rise. Almost any laser treatment of the material causes evaporation whose intensity varies between laser pulses of milli and nanoseconds (Fig. 11.1). Short high-power pulses are characterized by low evaporation and small amount of removed substances, while pulses of long duration and low power cause the removal of a large amount of material, which involves the formation of deep craters. The release of steam from the irradiated surface takes place at an almost sonic speed in the form of a plume directed from the surface. In addition to steam, this plume contains drops of liquid phase as well as solid particles removed from the bottom and sides of the crater"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001503_978-3-319-97304-3_34-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001503_978-3-319-97304-3_34-Figure1-1.png",
        "caption": "Fig. 1. The mobile robot environment containing one target and one closest obstacle.",
        "texts": [
            " In the states and actions definition, we care about only the approximate distance and relative directions between the robot and the target, and between the robot and the closest obstacle. At each time instant, the robot is considered as the center, and the environment around it is divided into four regions: R1, R2, R3 and R4. To avoid the situation that the obstacle and the target may locate at the same region which will confuse the robot, the angle between dr\u2212t (the line from robot to target) and dr\u2212o (the line from robot to obstacle) is considered, as shown in Fig. 1(a). The range of \u03b8 is [0, 2\u03c0], as we only need to know the range of the angle, thus, we divide the interval into eight angular regions, which is shown in Fig. 1(b). Assume that at time instant t, the robot\u2019s position is Pr = [Px Py]t, the target position is Pt = [Pt\u2212x Pt\u2212y]t, and the closest obstacle location is Po = [Po\u2212x Po\u2212y]t. Then we have \u03b8 = arcsin D dr\u2212o , (1) where D = |(Pt\u2212y \u2212 Py)Po\u2212x + (Px \u2212 Pt\u2212x)Po\u2212y + Pt\u2212xPy \u2212 PxPt\u2212y|\u221a (Pt\u2212x \u2212 Px)2 + (Pt\u2212y \u2212 Py)2 , (2) and dr\u2212o = \u221a (Po\u2212x \u2212 Px)2 + (Po\u2212y \u2212 Py)2. (3) The current angular region G containing \u03b8 is depicted as follows: G = \u23a7 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 G1, \u03b8 \u2208 [0, \u03c0/4) G2, \u03b8 \u2208 [\u03c0/4, \u03c0/2) G3, \u03b8 \u2208 [\u03c0/2, 3\u03c0/4) G4, \u03b8 \u2208 [3\u03c0/4, \u03c0) G5, \u03b8 \u2208 [\u03c0, 5\u03c0/4) G6, \u03b8 \u2208 [5\u03c0/4, 3\u03c0/2) G7, \u03b8 \u2208 [3\u03c0/2, 7\u03c0/4) G8, \u03b8 \u2208 [7\u03c0/4, 2\u03c0)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000053_1464419314566086-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000053_1464419314566086-Figure3-1.png",
        "caption": "Figure 3. Roller load analysis in a tilting inner ring SRB.",
        "texts": [
            "comDownloaded from Ge \u00bc 7:3242 10 4k 0:2743e R 1=3ex \u00f013\u00de The total normal elastic deformation between the roller and the raceways can be expressed as n \u00bc i \u00fe e \u00bc Gi \u00fe Ge\u00f0 \u00deQ2=3 \u00bc GnQ 2=3 \u00f014\u00de where Gn \u00bc 7:3242 10 4\u00f0k 0:2743i R 1=3ix \u00fe k 0:2743e R 1=3ex \u00de \u00f015\u00de Then, the contact load due to elastic deformation can be expressed as29 Q \u00bc Kn 1:5 n \u00f016\u00de In equation (16), Kn is the stiffness coefficient, which can be calculated by Kn \u00bc G 1:5n \u00bc 5:045 104 k 0:2743i R 1=3ix \u00fe k 0:2743e R 1=3ex 1:5 \u00f017\u00de The normal loads of the jth roller Q1 j and Q2 j illustrated in Figure 1 can be expressed as Q1 j \u00bc Kn 1:5 1 j \u00f018\u00de Q2 j \u00bc Kn 1:5 2 j \u00f019\u00de The normal elastic deformation of roller raceway 1 j and 2 j are expressed as33 1 j \u00bc r cos 1j Pd 2 \u00f020\u00de 2 j \u00bc r cos 2j Pd 2 \u00f021\u00de where r is the inner ring radial shift and Pd is the bearing radial clearance. The half of load zone angle L can be calculated by the following equation L \u00bc arccos Pd 2 r \u00f022\u00de Figure 3 shows the roller load analysis in an aligned SRB. When the inner ring has a tilting angle, the loading direction of the bearing will also have a tilting angle . So, along y0 y0 direction, a static equilibrium relation can be obtained Fr1 \u00fe Fr2 \u00bc Fr=cos \u00f023\u00de The sum of components of Q1 j and Q2 j in x direction should be equal to zero, i.e. Fx1 \u00fe Fx2 \u00bc 0 \u00f024\u00de where Fr1 \u00bc X j\u00bc l j\u00bc0 Q1 j cos cos\u00f0 1j \u00fe \u00de; Fr2 \u00bc X j\u00bc l j\u00bc0 Q2 j cos cos\u00f0 2j \u00fe \u00de Fx1 \u00bc Xj\u00bcZ j\u00bc1 Q1 j cos sin\u00f0 1j \u00fe \u00de; Fx2 \u00bc Xj\u00bcZ j\u00bc1 Q2 j cos sin\u00f0 2j \u00fe \u00de at Kungl Tekniska Hogskolan / Royal Institute of Technology on September 10, 2015pik"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000665_med.2013.6608746-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000665_med.2013.6608746-Figure1-1.png",
        "caption": "Fig. 1: UAV with coordinates definitions.",
        "texts": [
            " It is worth highlighting that the airplane has an asymmetrical airfoil with no ailerons. The lack of ailerons can be considered a peculiar and quite important feature of the considered platform, since it has determined a greater difficulty in controlling the flight of the UAV with respect to other airframe platforms that allow the use of ailerons as, for instance, in [8]. The considered UAV has conventional horizontal and vertical tails with rudder \u03b4r and elevator \u03b4e control surfaces as shown in Fig. 1. The fuselage was made up of a carbon stick and wings are put together forming a dihedral angle of 22\u00b0. Some UAV physical parameters are given in Table I. The standard 6-DOF equations of motion for a conventional aircraft are used for modeling and simulating the small-size UAV. Supposing the aircraft as a symmetric (with respect the x \u2013 z plane) rigid body and the Earth flat (the airplane is flying over a small area), the body-axes forces, moments, kinematics and navigation equations are described below [9]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001777_978-1-4419-8420-3-Figure2.8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001777_978-1-4419-8420-3-Figure2.8-1.png",
        "caption": "Fig. 2.8 Simulation environment. The 3-D model of the FAMPHER robot in the pipeline; red points denote guarding points",
        "texts": [
            "3\u2229; the CCD camera is Point Grey Research Flea2, which measures 330 points). The height information can also be obtained by laser scanners in [15, 30] or other range cameras (e.g., [26, 42]). The rotation of the sensor can be controlled by a small mechanical platform installed in the rear part of the robot; therefore, the scan can be conducted radially inside the pipe toward different directions. Several effective sensing platforms with similar mechanism have been developed [30]. We develop a simulated platform for testing our algorithm (see Fig. 2.8), which simulates the process of inspection. We apply our procedure on this platform using several complicated 3-D virtual pipelines. The simulated results are convincing and show the effective inspection on pipeline geometry. 32 2 Region-Guarding Problem in 3-D Areas If a hole appears on the pipeline, it can be identified online when the robot reaches the guarding point that covers this region and matches the captured range depth images with the stored templates. We simulate this on pipeline meshes M by randomly generating some missing regions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001914_ilt-09-2013-0101-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001914_ilt-09-2013-0101-Figure2-1.png",
        "caption": "Figure 2 Schematic view of test rig for effects of misalignment",
        "texts": [
            " The main feature of this test rig is the non-rotating shaft mounted on a rotary table that allows us to incline the shaft at any desired angle. Meanwhile, the bearing sleeve is restricted to move only along one direction on a long linear bushing. Consequently, the bearing sleeve is not perpendicular to the shaft, thus indicating that the bearing is misaligned. Thus, we can adjust the degree of misalignment of GFBs freely by turning the rotary table. The schematic view of the test rig is presented in Figure 2. The length-to-diameter ratio of the linear bushing is 90/20 mm, thus ensuring that the axial runout deviation of the rod is negligible. The minimum adjustment angle of the rotary table is 20 (0.0056\u00b0), which is small enough compared with the misalignment tolerance of GFBs. A photograph of a Generation I bump-type test foil bearing is displayed in Figure 3. We follow the manufacturing process described in DellaCorte et al. (2008) to obtain the test GFB. The measurements in the present study are conducted under room temperature; thus, we use a tin solder to fix the foil sheets to the sleeve for convenience"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002012_msec2014-4029-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002012_msec2014-4029-Figure1-1.png",
        "caption": "FIG 1: POWDER INJECTION LASER CLADDING TECHNIQUE [13]",
        "texts": [
            " Consequently, this work is focused on the development of 3-D coupled thermomechanical finite element modelling in ABAQUS\u00ae for laser cladding of CPM9V (crucible steel) on H13 tool steel. The addition of new elements to the substrate is simulated by using the element birth technique. The clad dilution, heat affected zone and the residual stresses have been predicted from the model and compared with the experimental results to ensure sound repair. In this work, the powder injection mode of laser cladding is considered where the powdered material is injected from a nozzle and is deposited over the base material. Figure 1 shows schematically description of the process. In this work the substrate employed is H13 tool steel and clad material is Vanadium steel, CPM9V. Chemical composition of H13 tool steel is provided in Table 1 and the chemical composition of CPM9V is provided in Table 2. Actual cladding has dimension of substrate as 125 mm\u00d7 105 mm\u00d7 15 mm over which clad was deposited at length of 4cm. Uniform laser heat source is employed with beam diameter of 3mm and power of 2000-3800 W. Powder particles were spherical in shape with size of 44-104 \u00b5m"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003511_i2mtc.2018.8409574-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003511_i2mtc.2018.8409574-Figure4-1.png",
        "caption": "Fig. 4. Faults in the inter-shaft bearing",
        "texts": [
            " The faults are introduced to the inter-shaft bearing under the running conditions of high-voltage-motor single rotation (HR), lowvoltage-motor single rotation (LR), and high-voltagemotor/low-voltage-motor relative rotation (HLR), respectively. In addition, a normal condition of a two motor relative rotation is tested. Therefore, it namely becomes a pattern recognition problem on 10 different faults. The rotation speed of the motors is 20 Hz. The fault grooves in the bearing are machined by an electric spark, as shown in Fig.4. The outer ring fault is displayed by taking one of roller elements because the outer race and the holder cannot be removed. The experiment was performed four times under the same conditions and each condition consisted of 10 seconds of data. The data was divided into 200 samples, in which each sample is a measured vibration signal consisting of 1200 sampling data points. Three-quarters of the data were randomly selected to serve as the training set and the remaining 1/4 was used as the test set, and the details of the samples are shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000465_s00170-013-4872-6-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000465_s00170-013-4872-6-Figure8-1.png",
        "caption": "Fig. 8 Displacement of cutter",
        "texts": [
            " Let P1 Q1x;Q1y;Q1z; i1; j1; k1 and P2 Qx2;Qy2;Qz2; i2; j2; k2 be any two neighboring feature points (including both the cutter tip center points and the corresponding posture vectors), linear interpolation is taken for the cutter tip center points as (9). A new cutter center point path is obtained, as shown in Fig. 7. The new cutter center point path passes all the original cutter center tip points. Qtx \u00bc Q1x \u00fe t Q2x Q1x\u00f0 \u00de Qty \u00bc Q1y \u00fe t Q2y Q1y Qtz \u00bc Q1z \u00fe t Q2z Q1z\u00f0 \u00de 8 < : t \u00bc 0; 1 N ; ; N 1 N \u00f09\u00de The original cutter posture vectors are corresponding to the original cutter center tip points one to one. The change of cutter posture vectors may cause cutter interference and overcut to the machining parts. Figure 8 shows the displacement of cutter. In order to prevent from the cutter interference and overcut, the original cutter posture vectors are kept corresponding to the original cutter center tip points. For the interpolated points between two neighboring cutter tip center points, linear interpolation is utilized to generate the corresponding cutter posture vectors. Let P1 Q1x;Q1y;Q1z; i1; j1; k1 and P2 Qx2;Qy2;Qz2; i2; j2; k2 be two cutter location data on the original path, the cutter posture vectors corresponding to the intermediate interpolated points can be computed by (10)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003887_j.ymssp.2019.04.039-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003887_j.ymssp.2019.04.039-Figure2-1.png",
        "caption": "Fig. 2. Wind turbine-gear-generator train.",
        "texts": [
            " Firstly, the lumped parameter model for the machine chain is built to compute the gear meshing forces [5,19]; secondly, the FEM (Finite Element Method) model is built to obtain the Frequency Response Functions (FRFs) of the transfer paths; thirdly, the discrete spacetime model is built based on the continuous space-time model to compute the response signal. The Lumped parameter model targets a wind turbine-gear-generator train comprised of one turbine, one gear transmission and one generator (cf. Fig. 2), the gear transmission of which contains three stages, i.e. Stage 1 planetary gears, Stage 2 planetary gears and Stage 3 parallel gears. The model includes 44 degrees of freedom (DOF), i.e. one rotational DOF for the turbine; one rotational DOF for the rotor of the generator; two translational DOFs and one rotational DOF for each of the six gears of Stage 1, the six gears of Stage 2, and the two gears of Stage 3. The model adopts the below assumptions. a) The gear bodies of the ring, planet and sun are assumed rigid"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001258_j.jmatprotec.2017.11.010-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001258_j.jmatprotec.2017.11.010-Figure3-1.png",
        "caption": "Fig. 3 Flow forming process schematics",
        "texts": [
            " Firstly, to analyze the flow forming properties of Inconel 718 from the point of view of future applications - mostly aerospace industry requirements. The second was to analyze the influence of total strain on the microstructure and mechanical properties of heat treated flow formed elements. 2. Experimental The studied material was Inconel 718 obtained according to AMS 5596. Metal sheets were initially pressed to form cylinders according to the geometry shown in Fig. 2. Afterwards they were subjected to flow forming (Fig. 3) with 4 different thickness reductions: 35%, 40%, 55% and 65%. It is worth noting that the geometry of the rollers is not standard. The general idea of the conducted project was to combine three metal working techniques, conventional metal spinning, shear and flow forming. A compromise was established between the roller geometry in order to avoid tool rearming of rollers. Additionally, the manuscript machine is a prototype model with built in laser that can be used for heating of the work piece"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000757_j.cirpj.2015.08.005-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000757_j.cirpj.2015.08.005-Figure1-1.png",
        "caption": "Fig. 1. Axial and lateral magnetic force direction in the mobile magnetic clamp.",
        "texts": [
            " The magnetic clamp [1] main function is to hold a panel while pockets are machined on its surface. At the same time since it must slide over the panel, friction forces are generated. Axial magnetic attraction force between three closely located pairs of cylindrical magnets provides the support force to keep the panel in position. Lateral attraction force between the cylindrical permanent magnet pulls the slave part of the clamp over the panel to keep it close to the master part of the clamp. Axial and lateral magnetic force directions are shown in Fig. 1. Vuc\u030ckovic\u0301 et al. [2] presented a semi-analytical approach for the determination of the magnetic levitation force between two laterally displaced cylindrical permanent magnets. Fictitious magnetization charges and the discretization technique were adopted for cylindrical magnets assuming similar magnetic material and uniform magnetization along their axes of symmetry (opposite direction). \u00a7 This work was supported by the project \u2018\u2018Machining of skin panels (Al or Al-Li)\u2019\u2019, CRIAQ MANU-412 \u2013 NSERC RDCPJ 411911 \u2013 10)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001361_j.matpr.2017.11.041-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001361_j.matpr.2017.11.041-Figure1-1.png",
        "caption": "Fig 1. Test fixture",
        "texts": [
            " Investigations were carried out using discs of various materials such as Copper, Aluminum and Brass which have different material properties. The experiments were conducted with discs of various thicknesses and at various speeds of discs. Table 1 indicates various properties of Copper, Aluminum and Brass materials. With the rapid expansion of high speed lines and highways around the world, there is clearly considerable potential for a wider application of frictionless braking. The investigations were carried out on a test fixture as shown in Fig 1. The base is made of Aluminum in which a DC motor of 10,000 rpm is mounted. Stainless steel shaft is coupled to motor on which various discs of Copper, Aluminum, and Brass are mounted. The speed of disc is controlled using dimmer stat. Permanent magnets are moved towards and away from rotating disc using stepper motors. The values of Magnetic flux density of magnets and speed of discs are acquired through Hall Effect sensor and optical speed sensor respectively which are mounted close to the disc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001250_j.matpr.2017.07.234-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001250_j.matpr.2017.07.234-Figure4-1.png",
        "caption": "Figure 4: Contour parallel strategy",
        "texts": [
            " The surface quality of the near-net shape is not of great importance in HLM as it subsequently undergoes 8840 Sajan Kapil / Materials Today: Proceedings 4 (2017) 8837\u20138847 a much more accurate finish-machining operation, but integrity of the interior is very important. In most cases, contour-parallel strategy is preferred for HLM because it can provide: Effective heat dissipation, Less number of arc switch on/offs, Reduced warpage as toolpath direction changes continuously, Better geometrical quality of object by following geometrical trend of boundaries. The contour parallel strategy for a random boundary is shown in Fig.4. Several deposition methods have been studied and in most of them the deposition has been started on a base metal plate which has to be removed after deposition. In this section a new integrated substrate method has been explained in which the substrate becomes a part of the component. 1. It has been found that if an integrated substrate is used which becomes part of the deposited component then it increases the strength of the part as the forged/rolled substrate is reinforced and it can be worked as skeleton for the component"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000651_1.3656586-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000651_1.3656586-Figure1-1.png",
        "caption": "Fig. 1 Th in fi lm between near ly paral lel surfaces",
        "texts": [
            " II Generalized Thin Film Lubrication Equations The theory of lubrication is based on the Navier-Stokes equation with simplifications made possible by the fact that viscous shear usually predominates over fluid mass inertia and that the fluid film thickness is always considerably smaller than the characteristic dimension of the bearing surface. These stipulations can be expressed as the following mathematical statements: \u2014 I pdn + di f Jo dp dx' pudn 1=0; d_ dn dp dti = 0 (1) (2) (3) Equation (1) expresses the fluid continuity law in the space be- tween two nearly parallel surfaces separating from each other by the gap h as shown in Fig. 1. Equations (2) and (3) constitute the simplified version of the Navier-Stokes equations by virtue of the assumptions f = 0(1); 1 Prepared under Contract Nonr-3730(00), Task NB 061-131. Jointly supported by the Department of Defense, Atomic Energy Commission, and the National Aeronautics and Space Administration. Administered by the Office of Naval Research, Department of the Navy. Reproduction in whole or in part is permitted for any purpose of the U. S. Government. Contributed by the Lubrication Division of T H E A M E R I C A N S O C I E T Y OP M E C H A N I C A L E N G I N E E R S and presented at the ASMEASLE Lubrication Conference, Pittsburgh, Pa"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003766_rcs.1983-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003766_rcs.1983-Figure2-1.png",
        "caption": "FIGURE 2 Schematic model of the one-DOF joint tendon-driven mechanism (subscripts 1 \u2212 4) and the grasper actuation mechanism (subscript g)",
        "texts": [
            " The four joint actuation wires pass through the holes from 1 to 4, which are arranged every 90\u25e6, respectively. The backbone is inserted through the center hole. The backbone enhances the joint stiffness and guides the grasper actuation wire to pass through the center of the joint. In the joint actuation mechanism, we adopted the two-DOF tendon-driven mechanism, which is driven by the four pneumatic cylinders. The cylinders (SMC, Corp, CJ2XB10-15) are mounted every 90\u25e6 to correspond the wire holes. Figure 2 shows a schematic model of one-DOF joint mechanism driven by the cylinder displacements X1 and X3, which are measured by the linear potentiometers (Alps Electric, Co, Ltd, RDC1010A12). On a plane orthogonal to this plane, the other one-DOF joint, which is driven by the cylinder displacements X2 and X4, is installed. These subscripts correspond the wire holes in Figure 1. Each cylinder rod is connected to the wire. These wires are guided by the pulleys, go through the sheath, and are attached to the tip disk of the joint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000652_lmag.2012.2214027-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000652_lmag.2012.2214027-Figure2-1.png",
        "caption": "Fig. 2. Magnetic potential contour plot with superimposed field lines (dashed) of a uniformly magnetized cube of volume 1 cm3 and saturation magnetization \u03bc0 M0 = 1 T whose center is located 1 cm below the interface with a soft magnetic half-space of infinite permeability (represented by the thick black line on top). The cube is rotated by (a) 0\u25e6, (b) 45\u25e6, (c) 90\u25e6 with respect to the surface normal. The potential contour labels are specified in ampere.",
        "texts": [
            " To calculate the resulting magnetic field according to [Beleggia 2003], we describe the magnet by its indicator function D(r) that equals to 1 inside the body and zero outside it, so that its uniform magnetization vector is M1 = M0 D(r)m1, where m1 = [mx , my, mz ] is a unit vector and M0 is the saturation magnetization of the material. As dictated by the symmetry rules just derived, the image magnet will have magnetization M2 = M0 D(x, y, \u2212z + 2z0)[\u2212mx , \u2212my, mz ]. The shape amplitude (Fourier transform of D(r)) of the image magnet is then D(kx , ky , \u2212kz)exp(\u22122i z0kz), where D(k) is the shape amplitude of the real magnet. The magnetic potential in momentum space is then the product between the dipole potential and D(k), and its real space expression is obtained by inverse Fourier transform. As an example, Fig. 2 shows a contour plot of the potential of a uniformly magnetized cube located below a soft magnetic half-space tilted with respect to the surface normal. The interaction between the magnet and its image counterpart is what determines the net force, which can be calculated explicitly. As illustrated in [Beleggia 2004], the interaction energy between two uniformly magnetized bodies can be computed from their shape amplitudes. In particular, the interaction energy between the magnet and its image is E = 1 2 Kd 4\u03c0 3 Re \u222b d3k |k|2 D(k)D\u2217(\u2212ks)(m \u00b7 k)(ms \u00b7 k)e2i z0kz (8) where Kd = \u03bc0 M2 0 /2 is the magnetostatic energy density (shape anisotropy constant), and the subscript s denotes the mirror symmetry operation that changes the sign of the first two components of a vector and leaves the third unaffected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001802_ijsurfse.2017.088969-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001802_ijsurfse.2017.088969-Figure2-1.png",
        "caption": "Figure 2 Schematic diagram of the wear scar on the fat specimen against a spherical counter face ball",
        "texts": [],
        "surrounding_texts": [
            "where"
        ]
    },
    {
        "image_filename": "designv11_13_0003564_s00170-018-2592-7-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003564_s00170-018-2592-7-Figure6-1.png",
        "caption": "Fig. 6 Calculation domain in simulation",
        "texts": [
            " (2) The pressure in each cell is iteratively adjusted through implicit solution to satisfy the continuity equation; the resulted changes of the velocities are added to the velocities computed in step (1). (3) The fluid configuration is updated by solving the VOF equation in Eq. (5). The welding process was implemented using a laser power of 7 kWwith the welding speed of 1m/min along the x-direction. The simulation process lasted 1.8 s for laser welding and 1.0 s for cooling. A fluid region of 50mm\u00d7 14mm\u00d7 30mmwith a void region of 5 mm height above it was adopted to represent the calculation domain as shown in Fig. 6. The size of the mesh cell was 0.3 mm, and the total number of the mesh cells was 502,776. In addition, the coupled equations are solved with a time step of around 1.0 \u00d7 10\u22126\u20131.0 \u00d7 10\u22125 s. The physical properties of 5083 aluminum alloy used in the simulation are shown in Table 2. The keyhole dynamics, the molten pool fluid flow, and the weld bead formation under varied ambient pressures were analyzed. The keyhole dynamics of the laser welding under varied ambient pressures are discussed in this section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001714_1.g003503-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001714_1.g003503-Figure2-1.png",
        "caption": "Fig. 2 Rotating LVLH frame used to describe the relative motion of spacecraft.",
        "texts": [
            " The second type of graph is a line graph (i.e., there is only one path connecting all the nodes in the network as shown in Fig. 1b). The corresponding adjacency matrices are shown in Eq. (4): A 2 664 0 1 1 1 0 1 1 1 0 3 775 a ; A 2 664 0 1 0 1 0 1 0 1 0 3 775 b (4) To formulate the consensus estimation system of spacecraft relativemotion, it is necessary to understand how tomodel themotion of one spacecraft (deputy) relative to another spacecraft (chief). For this purpose, the local-vertical-local-horizontal (LVLH) frame [21] (shown in Fig. 2) is used. In Fig. 2, x denotes the coordinate in the radial direction O\u0302r, y denotes the coordinate in the along-track direction O\u0302\u03b8, and z denotes the coordinate in the cross-track direction O\u0302h. Assuming the orbit of the chief to be circular, the angular velocity of the LVLH frame with respect to an inertial frame, expressed in LVLH coordinates, is given by \u03c9 0; 0; n T, where n \u03bc\u2215r3c p is the mean motion of the chief satellite, \u03bc denotes the Earth\u2019s gravitational parameter, and rc denotes the chief orbital radius"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000203_j.ijmecsci.2013.11.010-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000203_j.ijmecsci.2013.11.010-Figure2-1.png",
        "caption": "Fig. 2. Schematics of McKibben actuator. (a) Overall shape, (b) cross-section, (c) diagram of unfolded sleeve.",
        "texts": [
            " Tel.: \u00fe81 926735605; fax: \u00fe81 926735090. E-mail addresses: chendh3124@gmail.com (D.H. Chen), kuniharu@ip.kyusan-u.ac.jp (K. Ushijima). elastomeric bladder made of butyl rubber (Young's modulus Er\u00bc2.5 MPa, Poisson's ratio \u03bdr\u00bc0.48) and nylon sleeve (Young's modulus En\u00bc2.0 GPa, Poisson's ratio \u03bdn\u00bc0.34). In the present study, some experimental investigations have been conducted to obtain the static relationship between the actuator's length L and the injected air pressure P under some several loads. Fig. 2(a) shows the geometry of the McKibben actuator. The parameters L, D and \u03b3 shown in Fig. 2(a) represent the length, the diameter and the rotational angle of a thread in a sleeve, respectively. Fig. 2(b) shows the cross section of the McKibben actuator, and Fig. 2(c) shows a spread-out mesh sleeve where n rounds (in the figure, n\u00bc2) of m (in the figure, m\u00bc3) threads of length b are wrapped around the outer surface of the rubber tube. As shown in Fig. 2(a)\u2013(c), the angle \u03b3 changes as the actuator expands and contracts. Here, the length L and the diameter D can be expressed in terms of the angle \u03b3 given by the following: L\u00bc b cos \u03b3; D\u00bc b sin \u03b3 n\u03c0 : \u00f01\u00de Fig. 3 shows the experiment setup of the McKibben actuator. In the experiments, a spindle, which is used to apply the prescribed tension force F, is first attached to the lower end of a suspended actuator, as shown in Fig. 3, and compressed air is gradually injected into and vented from the rubber tube, and the change of the length L is measured",
            " (3) that the maximum value of the angle \u03b3 can be given theoretically by \u03b3max \u00bc cos 1 1= ffiffiffi 3 p ffi54:71. Also, Chou and Hannaford [9] proposed another formula for the relationship between the tension force F and the pressure difference P\u2032 by considering the effect of tube thickness and the sleeve in a McKibben-type actuator as follows: F \u00bc \u03c0P\u2032 4 b n\u03c0 2 \u00f03 cos 2 \u03b3 1\u00de \" # \u00fe\u03c0P\u2032tk b n\u03c0 2 sin \u03b3 1 sin \u03b3 tk ; \u00f04\u00de where the actuator's thickness tk is determined by the distance in the thickness direction between the centerline of the mesh sleeve and the inner surface of the rubber tube, as shown in Fig. 2(b). They reported that Eq. (4) can predict the experimental data more accurately than from Eq. (3). Figs. 4 and 5 show comparisons of the relationship between the applied pressure P\u2032 and the contraction ratio \u03b7 obtained by experiment and from the analytical equations (Eqs. (3) and (4)). It can be seen that there is a substantial gap between experimental data and their predictions using (3) and (4). The main reason of the difference can be assumed by the unconsidered energy losses. One of the major energy losses is the frictional resistance",
            " Chou and Hannaford [9] have conducted approximate analyses by simply assuming a constant frictional force of 72:5 N for threads made of nylon and 75 N for threads made of glass fiber, which is based on the experimental data. In order to evaluate the possible resistance precisely, the analytical modeling for each resistance should be developed. In the following, the mechanics of sleeves in the deformation of McKibben-type actuator is discussed, and the theoretical modeling of the possible major resistance in the actuator under both pressurization and decompression process is proposed and evaluated analytically. Firstly, the geometry of a rhomboidal shape ADBC for threads as shown in Fig. 2(a) is discussed. Here, the length for each side of the rhombus can be expressed in terms of l0 as follows: AD \u00bc jDBj \u00bc jBCj \u00bc jCAj \u00bc b 2mn \u00bc l0: \u00f05\u00de Based on the fact that the bending deformation of a thread is much larger than the axial deformation, it is assumed that the length l0 remains unchanged during the actuator's deformation. In the lengthening and shortening deformation behavior of a McKibbentype actuator, the angle \u03b3 changes as the threads mutually rotate around each other during the deformation",
            " Next, the force T applied to a single thread in a sleeve as shown in Fig. 6 is considered. Here, the force T is decomposed into the components T\u03b8 and Tx in the directions of \u03b8 (circumferential direction) and x (axial direction), respectively. Also, the stresses in x- and \u03b8-directions acting on the rubber are termed sx;rub and s\u03b8;rub, respectively. The static equilibrium of forces in two directions are given as follows: T\u03b8\u00fes\u03b8;rub tRl0 cos \u03b3 \u00bc P\u2032Rkl0 cos \u03b3; Tx\u00fesx;rub tRl0 sin \u03b3 \u00bc \u00f0F\u00fe\u03c0R2 kP\u2032\u00de l0 sin \u03b3 2\u03c0R ; 8>< >: \u00f06\u00de where tR is the tube thickness as shown in Fig. 2(b), and Rk is the radius subtracting the thickness as Rk \u00bc R tk . Since the Young's modulus for the rubber is much smaller than that for the thread, the stress components sx;rub and s\u03b8;rub in Eq. (6) can be negligible. Therefore, the forces T\u03b8 and Tx can be simplified as follows: T\u03b8 \u00bc \u03c0RRkP\u2032 m tan \u03b3 ; Tx \u00bc \u00f0F\u00fe\u03c0R2 kP\u2032\u00de 2m : \u00f07\u00de Finally, the force T acting along a thread can be written as follows: T \u00bc T\u03b8 sin \u03b3\u00feTx cos \u03b3 \u00bc \u03c0R2P\u2032 cos \u03b3 m 1 2 Rk R 2 \u00fe Rk R \u00fe F 2\u03c0R2P\u2032 \" # : \u00f08\u00de As can be imagined from Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002445_ijmmme.2016010103-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002445_ijmmme.2016010103-Figure3-1.png",
        "caption": "Figure 3. FEA meshed model of gearbox",
        "texts": [
            " The numerical simulation was performed for the transient structural, thermal analysis of gear train under the influence of load, rotational speed and convection heat transfer coefficient. Gear oil works as gearbox lubricant to cool the gears and transfer the heat through convective process. The overheating of gears reduces the efficiency and life span of gears. The transmission gearbox assembly replacement is costly, so this study focus the thermal prospectus of gearbox so that the life span of a transmission gear train can be increased. Ansys 14.5 (Ansys, 2013), FEA simulation works on meshing of objects. The meshed model of gear assembly is shown in figure 3. The gear assembly is meshed using 5, 75, 383 nodes and 3, 39,898 elements. The present research work provides a strong base for the thermal-mechanical analysis of gearbox to increase the performance by understanding the role of gear oil and loading parameters such as torque, rotational speed and average convection heat transfer coefficient. Transmission gearbox components were assigned materials for accurate analysis of temperature on gearbox surface (Qin-man, 2011). The teeth in contact causes thin film thickness lubrication, thus a thermo-elastohydrodynamic arrangement of lubrication is expected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002720_978-3-319-24055-8-Figure3.3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002720_978-3-319-24055-8-Figure3.3-1.png",
        "caption": "Fig. 3.3 An example of the initial dynamic simulation, to illustrate the level of model complexity used in selecting the concept for dynamic performance",
        "texts": [
            " Representative values for each were collected from specialists amongst the project consortium, based on their past experience. The representative unit excitation values were not the actual values, since the gears and motor were yet to be defined in sufficient detail for the actual excitation to be calculated. Nonetheless, they were sufficient since the project was not, at this stage, aiming to give an absolute prediction of the vibration. Rather, it was aiming to understand the system response, guide design improvement and compare one layout with another (Fig. 3.3). The complexity of this process needs to be highlighted. At this stage of the design process there is no housing, so no value of surface vibration or radiated noise can be calculated. The approach used here was to calculate the total sound power transmitted through the bearings and use this as ametric in guiding the concept selection (Fig. 3.4). 36 B. James et al. The analysis team needed to check that this approach was valid, so an existing EV driveline (in production, complete with housing) was subjected to the same analysis method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003144_15397734.2017.1358094-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003144_15397734.2017.1358094-Figure2-1.png",
        "caption": "Figure 2. Diagram of cable model in three-dimensional inclined cable.",
        "texts": [
            " At the end we determined the static and dynamic stiffness of the CDPRs with respect to the global stiffness matrix. The static stiffness analyses is very essential for effective design, optimization and performance evaluation of the CDPRs. In this section, we introduced cable model, optimization of each cable lengths and tensions, and three-dimensional static stiffness analyses of inclined cable. Assume an inclined static sagging cable with a uniform cross-section and weight per unit length that connects two points A and B, with Cartesian coordinates (x, y, z) as shown in Figure 2. The configuration and parameter of cable are shown in Figure 2. The well-known catenary equations for three-dimensional by considering mass and elasticity can be written based on (Riehl et al. 2009; Yuan, Courteille, and Deblaise 2015). In Figure 2, the steel cable has unstrained crosssection area A(m2), elastic modulus E(Pa), cable length Lo(m) and self-weight q(N/m). In static equilibrium of CDPR the cables are attached between two end-points A and B. All the coordinates are expressed in the local frame attached to the three-dimensional inclined cable. The parameter \ud835\udeabL is the deformation of the cable due to self-weight and external force exerted on point B. Point p is an arbitrary point along the strained cable. This variable p belongs to the range 0 p L  which represents the strained length of the cable segment as measured from the end- point B of the cable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003766_rcs.1983-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003766_rcs.1983-Figure1-1.png",
        "caption": "FIGURE 1 Pneumatically driven surgical instrument",
        "texts": [
            " For each force estimation, we have constructed driving systems, kinematic models, dynamic models, and a force estimator. Finally, the accuracy of force estimation is evaluated by calculating the mean absolute error (MAE) between the measured value and the estimated value. In this prototype, the desired MAE is 0.2 N or less. We show that the mechanical interference is negligibly small by comparing the MAEs, which are obtained by changing the experimental conditions such as joint angles and grasping forces. The developed pneumatically driven surgical instrument is shown in Figure 1. Its length is 470 mm, the diameter of the insert portion is 5 mm, and its weight is 350 g. A tip unit has a two-DOF disk joint and a grasper. A drive unit consists of five pneumatic cylinders for driving the tip unit via wires, five linear potentiometers for measuring the cylinder position, and wire guide parts. In this work, we estimated the translation force and grasping force using a dynamics model Z and measured the driving force of the actuators F. For the estimation, the joint mechanism needs to be stiff against compression and requires a guide for a grasper actuation wire to avoid interference between a joint and a grasper driving mechanism. Therefore, we adopted a joint consisting of 10 disks and a flexible backbone (Covidien, SILS Hand Instruments). Since the disks are made of stainless steel, it has sufficient stiffness against compression. The disks are alternately stacked at 90\u25e6, and the click prevents twisting of the joint. The backbone has elasticity and guides the grasper actuation wire. The lower part of Figure 1 shows the cross-section diagram of the joint. The disk has five holes. The four joint actuation wires pass through the holes from 1 to 4, which are arranged every 90\u25e6, respectively. The backbone is inserted through the center hole. The backbone enhances the joint stiffness and guides the grasper actuation wire to pass through the center of the joint. In the joint actuation mechanism, we adopted the two-DOF tendon-driven mechanism, which is driven by the four pneumatic cylinders. The cylinders (SMC, Corp, CJ2XB10-15) are mounted every 90\u25e6 to correspond the wire holes. Figure 2 shows a schematic model of one-DOF joint mechanism driven by the cylinder displacements X1 and X3, which are measured by the linear potentiometers (Alps Electric, Co, Ltd, RDC1010A12). On a plane orthogonal to this plane, the other one-DOF joint, which is driven by the cylinder displacements X2 and X4, is installed. These subscripts correspond the wire holes in Figure 1. Each cylinder rod is connected to the wire. These wires are guided by the pulleys, go through the sheath, and are attached to the tip disk of the joint. The joint is driven by the joint torque \ud835\udf0fy , which is generated by the cylinder forces. Fi represents the cylinder driving force (N), which is computed by multiplying the air pressure Pij (kPa) by the pressurized area of the cylinder Aij (mm2). ui is the input voltage (V) to the servo valve. Servo valves, which is five ports spool type (FESTO, Corp, MPYE-5-M5-010-B), air pressure sensors (SMC, Corp, PSE540-01), and other electric and pneumatic components are built in the control box"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003971_j.procir.2019.03.008-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003971_j.procir.2019.03.008-Figure5-1.png",
        "caption": "Fig. 5. Adaptative Mesh.",
        "texts": [
            " It was placed at a distance of 0.3mm from the bottom surface. The objective is to suite each layer with elements whose height would be equal to that of a layer. Another partition has been created according to a plan passing almost by the tip of the piece to ensure a better mesh. It is also necessary to shift this partition to obtain a mesh as structured as possible. Finally, for a better respect of the symmetry of the piece, a last partition cutting the piece in half was created. The final part meshed is shown in the figure 5. Over 120000 8-node linear brick elements C3D8R are used. This work aims to check the influence of orthotropic behavior on the numerical results and to check the assumption of isotropy made in the previous work [1,2]. For that reason, two mechanical behaviors were implemented during the simulations. Only the behavior in elastic deformation was studied. Future work will aim to model the fracture behavior. Isotropic constants (table 2) will be based on the filament properties given by the technical data sheet [10] and the average value of Young\u2019s modulus in the three directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003797_s00170-018-03239-z-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003797_s00170-018-03239-z-Figure2-1.png",
        "caption": "Fig. 2 Multiblade skiving tool",
        "texts": [
            " On the contrary, if the tool provides the incremental movement, the relation can be represented as follows: \u03c92 \u00bc zg zt \u03c91\u2212 2v0sin\u03b21 mnztcos\u03b22 \u00f03\u00de where \u03b22 is the helix angle of tool. The purpose of tool design is to achieve higher machining efficiency and accuracy, as well as longer tool life. But these requirements can hardly be met simultaneously in the conventional skiving tool which is similar to a shaper with only one cutting blade. Therefore, an efficient skiving tool with a tapered shape and multiple blades for cutting the tooth space is proposed to manufacture small gears in power skiving (see, Fig. 2). As shown in Fig. 3, in case of working an internal gear, the multiblade skiving tool allows four subblades to be involved in one pass cutting, while the traditional skiving tool requires four passes cutting to machine the total tooth flank. In the process of gear machining, as shown in Fig. 3a, the subblade at the bottom of skiving tool is the first blade cutting into workpiece, then the second blade, third blade, and fourth blade successively involving in the cutting process. After the tool finish one revolution, the first blade, second blade, third blade, and fourth blade generate the tooth spaces, respectively, in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003448_j.procir.2018.03.277-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003448_j.procir.2018.03.277-Figure7-1.png",
        "caption": "Fig. 7 Example of a drawing made by one of the students.",
        "texts": [
            " Beforehand it was considered unlikely that the threaded parts could be printed in plastic, but the design students did it anyway. They did however experience that the smallest plastic screws broke during assembly because they were not strong enough, and the threads on the spindle (see third part from the top in Fig. 6) needed a brush up to run smoothly. In addition to the 3D-printed vises that the students made in groups, each member also delivered a set of individually made 2D-drawings. They made drawings of ten parts plus a drawing of the assembly. An example of one of the drawings is shown in Fig. 7. It is not possible to know how the quality of the drawings would have been if the students had not taken the optional additional work with 3D-printing. However, it should be noted that the 14 students participating had from none (0!) up to max three comments on their set of 2Ddrawings, which is an indication of an increased level of understanding when comparing with those students not taking on this task. Thomas Haavi et al. / Procedia CIRP 70 (2018) 325\u2013330 329 4 Haavi et.al./ Procedia CIRP 00 (2018) 000\u2013000 The interview guide was developed based on the CDIO framework [8] and had this structure: \u2022 Project as a whole \u2022 Process \u2022 Process evaluation \u2022 Self-evaluation of own learning \u2022 Suggestions for changes 3",
            " Beforehand it was considered unlikely that the threaded parts could be printed in plastic, but the design students did it anyway. They did however experience that the smallest plastic screws broke during assembly because they were not strong enough, and the threads on the spindle (see third part from the top in Fig. 6) needed a brush up to run smoothly. In addition to the 3D-printed vises that the students made in groups, each member also delivered a set of individually made 2D-drawings. They made drawings of ten parts plus a drawing of the assembly. An example of one of the drawings is shown in Fig. 7. It is not possible to know how the quality of the drawings would have been if the students had not taken the optional additional work with 3D-printing. However, it should be noted that the 14 students participating had from none (0!) up to max three comments on their set of 2Ddrawings, which is an indication of an increased level of understanding when comparing with those students not taking on this task. Fig. 3 Photograph of vise no. 1. Made by mechanical engineer students. Fig. 4 Photograph of vise no"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000616_aim.2012.6265962-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000616_aim.2012.6265962-Figure7-1.png",
        "caption": "Fig. 7. Concept of Visual Lifting Stabilization",
        "texts": [
            " Motion of the robot is constrained by the followed equation that is given by differentiating C1 by time after the strike. \u2202C1 \u2202q q\u0307(t+1 ) = 0 (21) Then, the equation of matrix formation in the case of heel\u2019s bumping can be obtained as follows.\" M(q) \u2212(jT c \u2212 jT t K) \u2202C1 \u2202qT 0 # \u00bb q\u0307(t+1 ) Fim \u2013 = \u00bb M(q)q\u0307(t\u22121 ) 0 \u2013 (22) We can derive the dynamics regarding the toe\u2019s bumping based on the similar above process. IV. VISUAL LIFTING STABILIZATION This section propose a vision-feedback control for improving humanoid\u2019s standing/walking stability as shown in Fig. 7. We use a model-based matching method to measure pose of a static target object denoted by \u03c8(t) based on \u03a3H , which represents the robot\u2019s head. The desired relative pose of \u03a3R (reference target object\u2019s coordinate) and \u03a3H is predefined by Homogeneous Transformation as HdT R. The difference of the desired head pose \u03a3Hd and the current pose \u03a3H is denoted as HT Hd , it can be described by: HTHd( d(t), (t)) = HTR( (t)) \u00b7 HdTR \u22121 ( d(t)), (23) where, although HT R is calculated by \u03c8(t) that can measured by on-line visual pose estimation method [28], [29], we assume this parameter as being detected correctly in this paper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003140_ffe.12681-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003140_ffe.12681-Figure1-1.png",
        "caption": "FIGURE 1 Load analysis. A, The moving and dynamic load on gear",
        "texts": [
            " This work not only can be used to analyze the characteristic and behavior of the crack which can lead to the tooth breakage but also can provide a reference for investigating the mechanism of the tooth breakage in wind turbine in which the operating condition of the gears is more complex. With gear transmission, a tooth can be alternatively exchanged between single and double teeth\u2010meshings, which results in the tooth suffering the moving and dynamic load with rectangular fluctuation,2 as illustrated in Figure 1A. Under this load, the gear tooth root is subjected to the maximum bending moment with dynamic variation. Neglecting the effect of tooth profile, the maximum bending moment at tooth root can increase linearly at AB, BC, and CD parts, respectively, where AB and CD are the double meshing areas and BC is the single meshing area. Due to the high load in the single meshing area, the maximum bending moment can increase faster than that in the double meshing area. This maximum bending moment can lead to the upper area of the gear tooth root being subjected to a maximum tensile stress, which can easily induce the crack initiation and propagation at the stress concentration areas of the gear tooth root",
            " The variation of the maximum tensile stress is consistent with that of the bending moment at tooth root. Under the maximum tensile stress, the upper area of the gear tooth root can be deformed uniformly. Due to linear increase of the maximum tensile stress at both double and single meshing areas, the strain rate at the upper area of the gear tooth root is remained constant at each meshing area. During a whole meshing cycle, the strain rate at the upper area of the gear tooth root can be described as Figure 1B. As illustrated in Section 2.1, the upper area at tooth root is subjected to the maximum tensile stress with dynamic variation. Hence, the gear tooth breakage generally occurs at tooth root and becomes a typical failure. With an aim to identify the characteristics and behaviors of the crack initiation and propagation under dynamic maximum tensile stress at the spur gear tooth root, the MD simulation is carried out. In general, the essential component of the spur gear material is ferrite which has bcc crystal structure and a worse ability in dissolving carbon",
            " The applications of the central crack and the periodic boundaries are consistent well with our purpose to simulate the initiation and propagation of the crack at the inner of the bulk. However, the application of the periodic boundaries can lead to the self\u2010interaction of the crack, which is the limitation of our simulation. In order to reveal the behaviors of the crack initiation and propagation under dynamic maximum tensile stress at the spur gear tooth root, the MD model is subjected to the tensile load along the crystal direction of [1 0 1]. The strain rates induced by the tensile load are described as Figure 1B, and they are 6\u00d710\u22123 and 3\u00d710\u22123 per picosecond at the single and double meshing areas, respectively. According to the MD simulation theory, the interaction between 2 iron atoms generates the metal potential. In order to precisely describe the metal potential, Finnis and Sinclair25 proposed the EAM potential as U \u00bc \u2211 N i\u00bc1 \u2211 N j\u00bci\u00fe1 u rij \u00fe \u2211 N i\u00bc1 \u03b1i ffiffiffiffi \u03c1i p (1) where u(rij) was the pair potential between the ith and jth atomic nucleuses, \u03c1i was the local electronic charge density of the ith atom, \u03b1i was a coefficient related to the material characteristics, and N was the total atomic number",
            " To avoid the thermal effect, the initial velocities of the atoms are generated by using a random number generator with the specified seed at the specified temperature of 1 K. Before simulating, the model is fully equilibrated for 10 ps to stabilize the temperature under NVE ensemble which corresponds to a constant atomic number, constant volume, and constant energy. Besides, this work resets the temperature of the system by explicitly rescaling the atomic velocities to accelerate the equilibration. After equilibrating, the typical strain rate as indicated by Figure 1B is applied to simulate the crack initiation and propagation under NVE ensemble. The simulation takes 10 ps at each meshing area. The timesteps are 0.001 ps during both the equilibration and simulation processes. Due to the time\u2010consuming characteristic of the MD simulation, this work only considers a whole meshing cycle, but it can provide a reference to analyze the initiation of the fatigue crack. Extracting the central atomic plane along the crystal direction of 1 0 1 , the arrangement of the atoms on this plane at 16 ps is shown in Figure 3 by using the visual software OVITO"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000334_tac.2013.2274707-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000334_tac.2013.2274707-Figure2-1.png",
        "caption": "Fig. 2. (a) satisfies Assumption 1. (b) violates Assumption 1.",
        "texts": [
            "1 in [7], where it is shown that solvability of the invariance conditions is necessary for solvability of RCP by continuous state feedback. The goal of this section is to provide a second necessary condition. The result is presented for single-input systems only, as the multi-input result is still unknown. Note that the presented result requires no assumption on the placement of with respect to . The set , the intersection of a simplex and an affine space, is a polyhedron. Suppose that is its vertex set; thus, . Let . Similarly, suppose is a polytope with vertex set , and let . Define the cone Consider Fig. 2(a). Here is depicted as the shaded cone with apex at . The set is not only a polyhedron, but also a simplex. It is clear from the figure that is precisely . The next result says to solve RCP by continuous state feedback there must be a non-zero vector in that lies in . Theorem 1: Suppose . If by continuous state feedback, then . Proof: Suppose by way of contradiction that . Since , by Proposition 3.1 of [7], one can find a continuous state feedback such that . Let . If , then implies that . Thus, is an equilibrium of the closedloop system",
            " Then , so (A1) and (A2) are satisfied. Thus, if we restrict to single-input systems, then Assumption (A1)-(A3) include the generic case. Finally, (A4) is a simplifying assumption and is the only notable loss of generality. However, it is important to note that how intersects is determined by the choice of triangulation. If the designer chooses to disregard (A4), then either a trial and error style of synthesis must be used [8], [13] or other triangulations must be adopted [3]. Example 1: Assumption 1 is illustrated in Fig. 2(a). We observe that is a simplex intersecting the interior of , but it does not intersect the facet . Therefore (A1) and (A3) hold. Also is a simplex so (A2) holds. Fig. 2(b) illustrates a situation when Assumption 1 fails. In this case is a polytope, not a simplex. Moreover, it intersects along the segment . Thus, both (A1) and (A4) fail. The following basic properties of and are derived from the fact that they are formed as intersections of affine spaces and a simplex. Lemma 1: If , then . If , then . When conditions (A3)-(A4) hold for a simplex inside , certain restrictions on the index sets arise. The next result identifies those restrictions for a general simplex "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001266_1350650117743684-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001266_1350650117743684-Figure6-1.png",
        "caption": "Figure 6. CFD analysis results.",
        "texts": [
            " The genes of Pareto optimal solutions are more likely to be passed to offspring. After offspring generation is generated, the fitness of new population will be calculated again, and the Pareto optimal solutions of new population will be updated according to the fitness of new population. Then, a new offspring generation will be generated with same gene operation. The evolution process continuously circulates and the Pareto front updated generation by generation until the maximum number of generations Ni is reached. Figure 6(a) shows the pressure distribution of initial thrust air film. It can be seen that the pressure inside the chamber is relatively high and keeps steady. Then it decreases from the edge of chamber to the outlet of air film gradually. Figure 6(b) presents the velocity contour and streamlines of fluid field near the orifice in the radial direction along Plane A. (The location of Plane A is shown in Figure 2(a).) Vortices are found in the chamber. This phenomenon is coincident with the study of Chen et al.23 and Gao et al.25 Its LCC is 1414N, static stiffness is 161N/mm, and VFR is 7.41 l/ min. As the CFD calculation results are sensitive to the mesh resolution, mesh independence of computational mesh is discussed in this section. y\u00fe is the nondimensional distance, which is related to the resolution of computational mesh near the boundary layer",
            " Figure 7 illustrates the influence of orifice number on the performance of thrust air bearing. It indicates that the LCC, stiffness, and VFR of thrust air bearing all increase with the rise of orifice number. Besides, the air film thickness that is corresponding to maximum stiffness slightly increases with the growth of orifice number. However, when the orifice number increases from 10 to 12, the LCC only rises slightly, and the maximum stiffness nearly keeps steady. The trends of each curve can be explained as following. From Figure 6(a), it can be seen that the pressure inside the chamber is much higher than the pressure of air film. In the radial direction of air film, the pressure decreases gradually. While, in the circumferential direction of air film, the pressure first decreases and then increases between two orifices. With the increase of orifice number, the distance between two orifices in the circumference decreases, and the area of high pressure region increases consequently. Therefore, the LCC increases with the increase of orifice number"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003202_j.ijsolstr.2017.10.008-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003202_j.ijsolstr.2017.10.008-Figure2-1.png",
        "caption": "Fig. 2. Side view of Dome I.",
        "texts": [],
        "surrounding_texts": [
            "Contents lists available at ScienceDirect\nInternational Journal of Solids and Structures\njournal homepage: www.elsevier.com/locate/ijsolstr\nSnap-through of shallow reticulated domes under unilateral\ndisplacement control\nRaymond H. Plaut\nDepartment of Civil and Environmental Engineering, Virginia Tech, Blacksburg, VA 24061, USA\na r t i c l e i n f o\nArticle history: Revised 4 October 2017 Available online 14 October 2017\nKeywords: Reticulated domes Lattice domes Snap-through Displacement control Force control\na b s t r a c t\nSnap-through instability of shallow reticulated (lattice) domes subjected to a quasi-static downward displacement or force at a joint is analyzed. For the case of unilateral displacement control, the joint is pushed downward by an indentor until it snaps (jumps) to another equilibrium configuration, and then pushed further until another snap occurs, and so on. Under force control, the magnitude of the force is increased, and a different sequence of snaps (local and then global) is exhibited. Green\u2013Lagrange strain is assumed, as well as engineering strain for the smaller dome. The equilibrium equations are solved numerically using Mathematica. For force control, snaps may occur on the equilibrium path (force versus displacement at the force) at limit (maximum) points and at bifurcation points. For unilateral displacement control, snaps may occur at points where the associated downward force decreases to zero, at turning points (which have a vertical tangent), and at bifurcation points (if the bifurcating path moves downward at the loaded joint).\n\u00a9 2017 Elsevier Ltd. All rights reserved.\no r o (\nd c c p i f s p\nD d D o m l m\nc a S i\n1. Introduction\nSnap-through of two shallow single-layer reticulated domes (caps) is analyzed. The joints (nodes) are pinned, with the support joints being immovable, and the members are assumed to remain straight, so that the structures are modeled as trusses. The members are uniform and linearly elastic, and have the same constant axial stiffness. The effect of the self-weight of the members is neglected.\nThe trusses are loaded quasi-statically at a single joint, using either unilateral displacement control or, for comparison, typical force control. Here unilateral displacement control means that a device (indentor) pushes monotonically downward on a joint. It should not be confused with the term \u201cdisplacement control\u201d describing a technique used in following an equilibrium path (e.g., Thai and Kim, 2009 ). At the pushed joint, the structure is free to move downward from the indentor but not upward. Such loading was considered in Plaut (2015a,b ) and references cited in those papers. Previous investigations of the stability of reticulated domes have focused on force control, for which (in the present analysis) the downward force at a joint is increased monotonically.\nSnap-through refers to a sudden jump in position of one or more joints. The dynamics of snap-through is not considered here, and it is assumed that the structure is damped and moves from one equilibrium state to another. Here snap-through may be local, in which a portion of the dome moves downward discontinuously,\nE-mail address: rplaut@vt.edu\nhttps://doi.org/10.1016/j.ijsolstr.2017.10.008 0020-7683/\u00a9 2017 Elsevier Ltd. All rights reserved.\nr global, in which the entire dome jumps to an inverted configuation below the horizontal plane of the support joints. A sequence f local snaps may occur before global snap-through takes place e.g., Shi et al., 2015 ).\nThe force on the loaded joint is assumed to be positive if it acts ownward (as it must in this analysis). For unilateral displacement ontrol, snap-through occurs either when the associated force dereases to zero, or when the equilibrium path exhibits a turning oint (a point with a vertical slope, with the path then changng direction), or sometimes at a bifurcation point in a plane of orce versus downward displacement of the joint. For force control, nap-through occurs when the equilibrium path exhibits a limit\noint or a bifurcation point.\nThe two domes analyzed in this paper are denoted Dome I and ome II, respectively. Dome I has been called a star dome, star ome truss, trussed star-dome, and hexagonal star-shaped dome. ome II is an extension of Dome I. The equilibrium equations are btained from stationarity of the total energy, and are solved nu-\nerically using Mathematica. For force control, stability of equiibrium states is determined by checking if the energy has a local\ninimum.\nThe formulation and numerical results for Dome I loaded at the rown are presented in Section 2 . In Section 3 , Dome I is loaded at\nring joint. Dome II is analyzed next, with loading at the crown in\nection 4 and at an upper-ring joint in Section 5 . Finally, concludng remarks are given in Section 6 .",
            "2\n2\ni ( 8 p p j s \u2026\ns ( h t a i a\nc ( m e L\nTable 1 Initial joint coordinates for Dome I.\nJoint i x i y i z i\n1 0 0 8.216 2 25 0 6.216 3 12.5 12.5 \u221a 3 6.216 4 \u201312.5 12.5 \u221a 3 6.216 5 \u201325 0 6.216 6 \u201312.5 \u201312.5 \u221a 3 6.216 7 12.5 \u201312.5 \u221a 3 6.216 8 25 \u221a 3 25 0 9 0 50 0 10 \u201325 \u221a 3 25 0 11 \u201325 \u221a 3 \u201325 0 12 0 \u201350 0 13 25 \u221a 3 \u201325 0\nL a i R W\nH c t I l ( T\nf c f a c f A V Y\nb 1 s f\n2\ni a f a j s p\nh\nT i\nL\nT\nG\n. Dome I loaded at crown\n.1. Previous work with force at crown\nFor the first example, consider the shallow dome (truss) shown n Figs. 1 and 2 , denoted Dome I. It has 24 members, 13 joints joint 1 at the crown, ring joints 2\u20137, and immovable support joints\n\u201313), and 21 degrees of freedom. The origin of the axes is in the lane of the support joints, underneath joint 1, with the x axis assing through support joint 2, the y axis passing through support\noint 9, and the z axis upward. The plane z = 0 will be called the upport plane. The initial coordinates x i , y i , and z i of joints i = 1, 2,\n, 13 are listed in Table 1 .\nThe ring joints lie on a horizontal circle of radius 25, and the upport joints lie on a horizontal circle of radius 50. The crown joint 1) has initial height 8.216, and the ring joints have initial eight 6.216. Dome I has been called a \u201cbenchmark problem for esting the accuracy of different geometric and material nonlinear lgorithms\u201d ( Yu et al., 2013 ). Dimensions of centimeters are listed n some previous studies of this dome, but any unit could be used nd no units are listed here.\nPapers analyzing this dome subjected to a vertical force at the rown (joint 1) include Almatian (2013), Ander and Samuelsson 1999), Blandford (1996a,b ), Greco et al. (2006), Hangai and Kawa-\nata (1972a,b,1973 ), Hartono (1997), Hill et al. (1989), Jagannathan t al. (1975a,b ), Koohestani (2013), Krishnamoorthy et al. (1996), ee and Han (2012), Leu and Yang (1990), Ligar\u00f2 and Valvo (1999),\nu et al. (2009), O\u00f1ate and Matias (1996), Papadrakakis (1981), Pardiso and Tempesta (1980), Paradiso et al. (1979), Ramesh and Krshnamoorthy (1993, 1994) , Rezaiee-Pajand and Estiri (2016a,b,c) ,\nothert et al. (1981), Tanaka et al. (1985), Thai and Kim (2009),\nriggers et al. (1988), Yang et al. (2007) , and Yu et al. (2013) .\nIn Abatan and Holzer (1978), Holzer et al. (1980), Watson and olzer (1983) , and Watson et al. (1983) , the initial heights of the\nrown and ring joints, respectively, are 8.161 and 6.161, and all he joints lie on a spherical surface of radius 157.25. For Dome , the joints lie close to a spherical surface. Domes with simiar topology but other dimensions were studied in Mahdavi et al. 2015), Saffari and Mansouri (2011), Saffari et al. (2008, 2014) , and oklu et al. (2015) .\nThe references listed in the previous two paragraphs consider orce control in which the vertical force at the crown was inreased or decreased along the equilibrium path. (In some cases, orces were applied to multiple joints and increased proportionlly.) Most the studies determine the first critical force for (loal) snap-through. The following compute the equilibrium path ar enough to include large displacements and several snaps:\nnder and Samuelsson (1999), Koohestani (2013), Ligar\u00f2 and alvo (1999), Watson et al. (1983), Wriggers et al. (1988) , and u et al. (2013) .\nFor Dome I, new aspects here include the loading of joint 2 y itself, and unilateral displacement control for loading of joint\nand, separately, loading of joint 2. Large displacements are conidered, and results are presented for Green\u2013Lagrange strain and or engineering strain.\n.2. Formulation\nAll members have the same constant axial stiffness EA , where E s the modulus of elasticity and A is the cross-sectional area. Either\ndownward force P 1 acts at joint 1 (the crown) or a downward orce P 2 acts at joint 2. (For unilateral displacement control, these re the forces associated with the indentor.) The displacements of oint i in the x , y , and z directions are denoted u i , v i , and w i , repectively, and the height of joint i above the horizontal xy support\nlane is\ni = z i + w i (1)\nhe initial length of the member (bar) ij connecting joints i and j\ns L ij , with its square given by\n2 i j = ( x i \u2212 x j )2 + ( y i \u2212 y j )2 + ( z i \u2212 z j )2 (2)\nhe deformed length of member ij is G ij , whose square is\n2 i j = ( x i + u i \u2212 x j \u2212 u j )2 + ( y i + v i \u2212 y j \u2212 v j )2 + ( h i \u2212 h j )2 (3)",
            "V\nIf engineering strain is assumed, the strain in member ij and the corresponding strain energy S ij are\n\u03b5 i j =\nG i j \u2212 L i j\nL i j\n, S i j =\nEA ( G i j \u2212 L i j )2\n2 L i j\n(4)\nIf Green\u2013Lagrange strain is assumed, the strain and strain energy in member ij are ( Ligar\u00f2 and Valvo, 1999 )\n\u03b5 i j =\nG 2 i j \u2212 L 2 i j\n2 L 2 i j\n, S i j =\nEA ( G\n2 i j \u2212 L 2 i j\n)2\n8 L 3 i j\n(5)\nThe total potential V is given by\n= S + P k w k (6) where S is the sum of the S ij and k = 1 or 2. The equilibrium equations are obtained from the stationarity conditions\n\u2202V\n\u2202 u i\n= 0 , \u2202V\n\u2202 v i = 0 ,\n\u2202V\n\u2202 w i\n= 0 (7)\nwhere i = 1, 2, \u2026, 7 for Dome I.\nThe analysis is conducted in terms of the nondimensional quan-\ntities\np 1 =\n10 4 P 1 EA , p 2 = 10 4 P 2 EA , \u03b41 = \u2212w 1 z 1 , \u03b42 = \u2212w 2 z 2 (8)\nwhere \u03b41 and \u03b42 are the nondimensional downward displacements of joints 1 and 2, respectively (e.g., \u03b41 = 1 when loaded joint 1 reaches the support plane, and \u03b42 = 1 when loaded joint 2 reaches the support plane).\nNumerical solutions of Eqs. (7) in nondimensional form are obtained with the use of the subroutine FindRoot in Mathematica. The force ( p 1 or p 2 ) is specified and initial guesses are chosen for the displacements u i , v i , and w i in Eqs. (7) . Different sets of guesses may lead to different equilibrium configurations. Segments of the equilibrium path are followed manually. For the case of force control, the stability of an equilibrium state is determined by applying the Mathematica subroutine FindMinimum to see if V has a local minimum there.\nFor loading at the crown, the equilibrium path is plotted as nondimensional force p 1 versus nondimensional downward crown displacement \u03b41 with p 1 \u2265 0 and \u03b41 \u2265 0. Segments of interest are presented in Fig. 3 for the ranges 0 \u2264 \u03b41 \u2264 2.5 and 0 \u2264 p 1 \u2264 100. Solid curves correspond to Green\u2013Lagrange strain and dashed curves to engineering strain. The dashed curves are indistinguishable from the solid curves except in a neighborhood of point E."
        ]
    },
    {
        "image_filename": "designv11_13_0002861_s12206-016-1035-3-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002861_s12206-016-1035-3-Figure1-1.png",
        "caption": "Fig. 1. (a) The target quad-rotor platform; (b) general configuration.",
        "texts": [
            " The quad-rotor, which is one of the multirotor configurations, has four motors fixed at four corners of its body. The most general type of a small quad-rotor was chosen as the target platform, which is shown in the Figs. 1(a) and (b). Further descriptions for 6-Degree of freedom (6- DOF) equation of motion is not required to be presented from this paper because it is already widely developed by many researchers [8, 11, 12], but only a few elements to describe the current application will be treated. It is shown in Fig. 1(b) the configuration of the motor thrusts 1 4~T T , torques 1 4~Q Q , and symmetric moment arm l . The body fixed reference frame B: ( bO , bx , by , bz ) and the local level inertial reference frame W: (O, x, y, z) are also shown where the bx axis points towards the middle of motor #1 and #4. It is also possible to use the bx axis to point toward the motor #1, but there is no fundamental difference, and it is freely switchable. The quad rotor platform used in this paper has 0.295 m length for the moment arm l , and takeoff-weight of 3",
            " The setpoint 1r is given for the primary measured variable 1y , and the primary loop control 1C yields the set-point 2r for the secondary measured variable 2y . It is known as a fact that the cascade control improves the set-point response because the secondary loop is operating with faster dynamics and higher gain which also leads to improved load disturbance [10]. Using the secondary measured variable as angular velocity, the cascaded attitude control \u2013 that is, Euler angle control - can be achieved without any additional sensor or calculation from the view point of a UAV. For such a small quad-rotor UAV as it was shown from Fig. 1, where the poorly streamlined structure of a UAV with extruded payloads increases the suscepti- bility to gust, the cascaded attitude control will provide a better performance [13-15]. However, when compared with a more straight-forward single-loop control, the major drawback of the cascade control systems is the complexity created by multiple control loops. If the primary controller 1C is to be designed, the process model is not simply 1 2G G but has a different form as: 2 2 1 2 2 1 C G GG C G = + (3) where the G is introduced to indicate the process model for primary controller, and it is the main reason for the increased complexity",
            " Therefore, the only problem left for the cascade control is the safety issue during the identification of the fast 2G dynamics. The safety issue depends on the stability of a target UAV platform. Generally, a small helicopter with a stabilizer bar or stabilizer paddle does not need any consideration but it only requires a skilled pilot who can manually fly the helicopter. However, a small quad-rotor platform can only reveal the second-order time-delay 2t with additional stability augmentation system simply because it is too fast. From the small quad-rotor shown in Fig. 1, 10 % of 2r input in roll or pitch axis yields 200~250\u00b0/s/s angular acceleration because the moment of inertia is very small. A systematic and analytic process to implement the cascaded attitude control is thus proposed as summarized with a process diagram in Fig. 4. Each blocks have the name of the step on the top side and key notes on the bottom side. The first problem of the safety issue is solved by having P control in the secondary control 2C , and an automated frequency sweep input. The first process, the \u2018auto-commanded test for secondary loop\u2019, initiates a frequency sweep with 1C control being turned off which is only possible in a very short time",
            ", Survey on attitude and heading reference systems for remotely piloted aircraft systems, 2014 International Conference on Unmanned Aircraft Systems (ICUAS), IEEE, Orlando, Florida, USA (2014) 876-884. [40] J. B. Song, Y. S. Byun, J. Kim and B. S. Kang, Guidance and control of a scaled-down quad tilt prop PAV, Journal of Mechanical Science Technology, 29 (2) (2014) 807-825. Appendix A.1 The choice of control signals and the distributing matrix for quad rotor It is followed a set of conventional control input terms widely used in aircraft control, i.e., the aileron, elevator, rudder, and throttle. Consider a quad-rotor represented in Fig. 1 that has four motors equivalently placed in distance l . Since the aileron is a control input to create a rolling moment, aileron (AIL) is defined from thrust ( iT , i = 1, 2, 3, 4) as : ( )3 4 1 2AIL a T T T T l= + - - \u00b4 (A.1) which is proportional to rolling moment. Similarly, the elevator (ELE) , rudder (RUD) and throttle (THR) are defined from thrust and torque ( iQ , i = 1, 2, 3, 4) as: ( )1 4 2 3ELE a T T T T l= + - - \u00b4 , (A.2) ( )2 4 1 3 1 4 2 3( )RUD a Q Q Q Q b T T T T= + - - = + - - (A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002679_0309324715614194-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002679_0309324715614194-Figure3-1.png",
        "caption": "Figure 3. Contact pressure distributions for (a) a single asperity and (b) an asperity of the wavy surface.",
        "texts": [
            "13,16 The calculations were carried out to compare the results for the model of one asperity and wavy surface, and also to analyze the influence of the surface geometry, adhesion, and sliding velocity on the contact pressure distribution, areas of contact and adhesive interaction, and deformation component of the friction force. For the calculations, the radius R of a separate asperity is taken to be equal to the radius of a tip of the wavy surface asperity R= l2/(2p2H). The results obtained depend on the following dimensionless parameters\u2014 viscosity parameter d=(Te/Ts), adhesion parameter l= p0(H/Eg) 1/3, velocity parameter k=(R/VTe), and the waviness parameter l/h for the periodic surface. In Figure 3, the distributions of the dimensionless contact pressure p(x, y)R/g are presented for one asperity (Figure 3(a)) and for the wavy surface (Figure 3(b)) as the same dimensionless normal load is applied to an asperity P/(Rg)=400. The distribution graph for the wavy surface (Figure 3(b)) is presented on the half of a periodicity cell (x=l) 2 \u00bd 0:5; 0:5 , (y=l) 2 \u00bd 0:5; 0 ; the graph for the separate asperity (Figure 3(a)) is presented on the same domain for the convenience of comparison. The values of the parameters used for calculation are d=10, l=1, k=0.4, l/h=4. The results indicate that viscoelastic properties of the foundation lead to the nonsymmetric pressure distribution (shifted in the direction of sliding). Adhesion leads to the negative constant pressure p(x, y)=2p0 acting in an area outside the contact region. For the case of wavy surface, the maximum contact pressure is smaller, while areas of contact and adhesive interaction are wider than for the separate asperity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002132_j.triboint.2014.09.007-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002132_j.triboint.2014.09.007-Figure2-1.png",
        "caption": "Fig. 2. Illustration of the reducing (or substructuring) technique applied to the connecting-rod big-end bearing considered in the present work.",
        "texts": [
            " Since the sole outer forces acting on the bearing structure are the hydrodynamic resulting ones, the elastic distortions of the bearing surface are here calculated using the reducing system approach. In this case, the bearing displacements and the hydrodynamic pressures are directly related by a compliance matrix. Such compliance matrix is computed in a pre-processing step using an external software (ABAQUS software was used in this work) by applying an appropriate substructuring technique to the complete 3D linear-elastic finite element model of the entire bearing structure (see Fig. 2). According to the aforementioned descriptions, the following vector\u2013matrix expression is used to represent the calculation of the bearing displacements from a given hydrodynamic pressure distribution \u03b4b S \u00bc C\u00bd SpH S \u00f04\u00de where C\u00bd is the compliance matrix and \u03b4b the vector of structural displacements. Throughout this work, lower-case and upper-case variables in bold indicate vector and matrix quantities, respectively. The superscript S denotes that all quantities are evaluated in the structural reduced mesh, described in Section 2",
            " Moreover, in order to allow a smooth change in the bearing geometry, the simulation of each load case is computed in 5 successive steps with a progressive variation of the external load until the aimed value is attained. Algorithms 3\u20135 illustrate the complete algorithms for the EHL solution of statically-loaded journal bearings for the different partitioned methods described in the previous section. In order to evaluate the performance of the proposed coupling algorithms presented in the previous sections, the connecting-rod bigend bearing illustrated in Fig. 2 is considered. For the definition of the reduced system and subsequent calculations of the compliance matrix, the degrees of freedom of the nodes located on a plane near to the main bearing are fixed in the entire FEM model. The imposition of such boundary conditions is required to suppress the free-free rigid body motion of the structure. However, the corresponding clamped condition tends to yield large, unrealistic bending deformations in the rod, which in turn have to be removed to properly evaluate the local influence of the structural distortions on the lubricant film thickness"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002340_1350650114562485-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002340_1350650114562485-Figure5-1.png",
        "caption": "Figure 5. The assumed variations of loads.",
        "texts": [
            " The bearings are located at the ends of the shaft, and the diameters at particular sections correspond to theoretical outline built according to the rule of equal flexural strength. The bearing located on the left end of the shaft is marked as \u2018\u2018A\u2019\u2019 and the one on the right end as \u2018\u2018B\u2019\u2019. The dimensions of the model shaft are: x2\u00bc 16mm, x3\u00bc 50mm, x4\u00bc 100mm, x5\u00bc 150mm, x6\u00bc 184mm, x7\u00bc 200mm, d1\u00bc 30mm, d2\u00bc 35mm, d3\u00bc 40mm, d4\u00bc 40mm, d5\u00bc 35mm, d6\u00bc 30mm. Nodal points in bearings (places of reaction concentration) are specified with coordinates: xA\u00bc 27mm, xB\u00bc 173mm. The assumed model of the shaft has been a subject to calculations for two different loads presented in Figure 5. In the first variation it is assumed that the load is applied on both sides of one gearwheel, placed at the distance xL from the beginning part of the shaft. Circumferential forces Fc are oriented in accordance with axis \u2018\u2018y\u2019\u2019, radial forces Fp in accordance with axis \u2018\u2018z\u2019\u2019 and axial forces Fx in accordance with axis \u2018\u2018x\u2019\u2019. In the second variation, the loads are located on two gearwheels located at distances xL1 and xL2 from the beginning part of the shaft, when they are located in points determined by angles b1\u00bc 90 and b2\u00bc 180 . The directions of loads are shown in Figure 5. In both variations, the axis forces sum up. at Gebze Yuksek Teknoloji Enstitu on December 19, 2014pij.sagepub.comDownloaded from The locations of planes have been assumed in determined relations up to the length of shaft Lw, equal to the dimension x7, and they are: \u2013 for variant I of the load: xL\u00bc 0.6 Lw\u00bc 120mm, \u2013 for variant II of the load: xL1\u00bc 0.4 Lw\u00bc 80mm and xL2\u00bc 0.6 Lw\u00bc 120mm. It has been assumed that loads in both the load points presented in Figure 5 are identical (Fc1\u00bcFc2, Fp1\u00bcFp2, Fx1\u00bcFx2). Thanks to the above, the torques affecting the shaft is balanced. The values of loads have been assumed in specific relations to basic dynamic load rating of bearings C. It has been established that the circumferential force on the putative gearwheel Fc1 will represent the level 0,1C. Provided that pressure angle of the gearing for this gearwheel equals approx. 20 , the radial force Fp has been determined as about 0.364 of the circumferential force. The axial force Fx has been assumed in five values in the following established relations to the peripheral force: 0, 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000842_tmag.2015.2480546-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000842_tmag.2015.2480546-Figure3-1.png",
        "caption": "Fig. 3. (a) Field winding scheme allowing imposing short-circuited parts. (b) Tooth winded sensor.",
        "texts": [
            " As the commercial frequency in Brazil is 60 Hz, the 50 Hz operation was chosen to segregate phenomena produced by the generator operation itself and those from the surrounding stray fields noise. Fig. 2 shows the 2-D machine calculation domain with the flux distribution at no-load during healthy operation. 0018-9464 \u00a9 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. The short-circuited turns in the generator rotor field winding can be experimentally imposed by applying the dc voltage only partially on the winding, as shown in Fig. 3(a). For instance, if we wish to simulate a short circuit between the points J and J2, the dc voltage is applied only between the points J2 and K. For analyzing the magnetic flux behavior, an induced voltage sensor is winded on a stator tooth, as shown in Fig. 3(b). In the simulation, the voltage induced in this sensor is calculated and compared with the experimental results. Simulation, as well as experimental waveforms, are acquired by means of discrete quantities (or sampled) as regularly performed in such simulations. The obtained amplitudes are related to pre-set time instants t1, t2, t3 . . . t f . One of the most usual ways to transform a discrete quantity in time domain to the frequency domain is the application of the FFT algorithm to calculate the Fourier discrete transformation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002012_msec2014-4029-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002012_msec2014-4029-Figure12-1.png",
        "caption": "FIG 12: LONGITUDINAL STRESS AT TIME (a) t= 3sec, (b) t= 10 sec",
        "texts": [
            " In laser cladding during a short time interval, the substrate is heated up to the melting point, and a molten layer of the clad material is deposited on it. Thereafter, the heat is conducted into the substrate and the clad solidifies and starts shrinking due to thermal contraction, whereas the substrate first expands and later contracts according to the local thermal cycle. This leads to the development of residual stresses. Figures 12 and 13 show the longitudinal stresses (along the scanning direction) variation with time for Case 2. Figure12 shows the longitudinal stress evolution during the time when the work-piece is cooling back to the room temperature. Figure 13 shows the longitudinal stress distribution after the work-piece has attained room Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 7 Copyright \u00a9 2014 by ASME temperature after sufficient interval of time. Figure12 indicates compressive nature of longitudinal stress in the clad region followed by tensile nature near the interface which subsequently changes back to compressive in the substrate. The contour plots in Figures 13 (a) and (b) show the similar trend, albeit, with slightly higher compressive stresses in the clad after complete cooling. The longitudinal stress in the interface region is tensile in nature which can be due to the fact that the interface region with small cross-section and comparatively higher thermal gradient than the underlying substrate material is being subjected to tensile load as it tries to cool down and attain steady state"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002066_2013.40949-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002066_2013.40949-Figure1-1.png",
        "caption": "FIG. 1 Cutaway view of a conventional tire.",
        "texts": [
            " At higher slips the advantage for radialply tires tends to drop off, but in the working range they show a definite traction advantage in most soil condi tions, but not in mud. There are other advantages too, but also some real disadvantages. Since the paper by R. L. Wann is concerned with the durability characteristics of radialply tires, this paper will concentrate on the operating characteristics, good or poor, of radial-ply tractor tires. The differences in these characteristics is apparently due to variations in the con struction of the two types of tires. Fig. 1 shows a cutaway view of a conventional tire. It has four plies of cord fabric, each ply cut on the bias and each ply opposite in cord angle to the adjacent plies. The ply cords ex tend from bead to bead in a crisscross fashion to provide a strong, sound, long-life tire. This basic construction is used by tire manufacturers the world over. There may be two, four, six, eight or more plies, and in some cases a multiply bias-cut breaker in the crown area of the tire either on top of the plies or buried between plies to pro vide extra bruise strength where needed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003482_s00170-018-2324-z-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003482_s00170-018-2324-z-Figure4-1.png",
        "caption": "Fig. 4 Cutting process at different times (rake angle and clearance angle are 15\u00b0 and 20\u00b0). a Beginning. b 0.01s. c 0.02s. d End",
        "texts": [
            " The advantage of the modified Coulomb friction odel is that the solution can determine friction state automatically according to the contact stress value in the cutting simulation process. In the oblique cutting simulation of TC21 alloy, the analysis step (Dynamic Temperature-displacement, Explicit) in the software ABAQUS 6.12 has been employed. This analysis step is used to indicate that a dynamic coupled thermal- stress analysis is to be performed using explicit integration. The cutting simulation of TC21 alloy is shown in Fig. 4. The chip formation and the distribution of stress at different cutting times were obtained through the simulation. As shown in Fig. 4, the primary shear band of chip can be observed clearly. The maximum stress was 1324 MPa and located in the primary shear band. The temperature distribution between the workpiece and tool in the cutting of TC21 alloy is shown in Fig. 5. Figure 5 shows that the chip and tool maximum temperature are 116.6 and 228.7 \u00b0C. And the region of the maximum temperature at tool and chip is in the contact surface between tool and chip and is 0.33 mm from the tip the region. Figures 6 and 7 are the distribution of residual stress in the cut layer of TC21 alloy after oblique cutting"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000650_13552541211193458-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000650_13552541211193458-Figure15-1.png",
        "caption": "Figure 15 3D metal samples of Cu-based powder mixture by DMLF",
        "texts": [
            " Top surface quality research for direct metal laser fabrication Jialin Yang, Hongwu Ouyang, Chao Xu and Yang Wang Volume 18 \u00b7 Number 1 \u00b7 2012 \u00b7 4\u201315 Above results show that metal substrate, powder composition, powder morphology and the balancing of laser power and scanning speed play crucially important roles in controlling TSQ of unit layer during DMLF of Cu-based powder mixture. Rapid manufacturing of 3D Cu-based metal samples by DMLF could be realized if spherical gas-atomized copper powder and suitable composition are chosen in powder mixture. For GA-Cu powder mixture (GA-Cu powder \u00fe CuSn20 powder \u00fe Cu3P powder), laser power and scanning speed should be selected in the range of 400- 600W and 0.05-0.20m/s, respectively, in order to obtain good TSQ. Figure 15 shows Cu-alloy samples prepared by DMLF using a laser power of 480W and a scanning speed of 0.15m/s. Figure 15(a) shows a sample with simple structure. Figure 15(b) shows a CAD model with four inner helix tunnels, which cannot be manufactured at all by means of traditional manufacturing technologies. But for DMLF, this structure can be easily and rapidly manufactured without the use of any tools or moulds, as shown in Figure 15(c). . TSQ is an exclusive technical term for DMLF compared with traditional manufacturing methods, which is very important for the process stability and part performances in DMLF process. It could be defined as the surface morphology in laser scanning plane, which should mainly include three requirements: flatness, compactness and cleanliness, and could be evaluated in macro and micro scopes. The good TSQ includes these characteristics: in the macroscopic view, a smooth, flat surface with no balling features, local peaks that may exceed the layer thickness or crack formation in the laser-scanned area, and the height of the inner of sintering surface being consistent with that of the boundary"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003607_j.procir.2018.08.113-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003607_j.procir.2018.08.113-Figure1-1.png",
        "caption": "Fig. 1. Additive manufacturing and forming of hybrid parts with an additively manufactured element in the a) tension zone and b) pressure zone",
        "texts": [
            " For a proper design of hybrid parts, the characterization of the bonding strength is essential. The interaction between subsequent process steps of different process routes and the influence on the shear bonding strength is already investigated in [5]. Furthermore, the stress state in the forming step could influence the bonding strength. Hence in this work, the influence of tension and pressure stress at the AE on the bonding strength is investigated. Therefore, the AE is placed on different sides of the bending specimen (Fig. 1). il l li t . i ir t. r ce ia I ( ) .else ier.c /l cate/ r ce ia - t rs. lis ls i r t . is is ss rti l r t - - li s ( tt :// r ti s. r /li s s/ - - / . /) eer-re ie er res si ilit f t e a erisc es aserze tr . t t i l i , , , , , , , , , , aI stit te f f ct ri ec l y ( ), rie ric - lex er- iversit t rl e - r er , erl str e , rl e , er y bI stit te f t ic ec l ies ( ), rie ric - lex er- iversit t rl e - r er , r - se- tr e - , rl e , er y c rl e r te c l i v ce tic l ec l ies ( ), l r tr e , rl e , er y rres i a t r",
            " i t i r , t i fl f t i r r tr t t t i tr t i i ti t . r f r , t i l iff r t i f t i i ( i . ). th CIRP onference on Photonic l i [ ] Thomas Papke et al. / Procedia CIRP 74 (2018) 290\u2013294 2912 Author name / Procedia CIRP 00 (2018) 000\u2013000 Nomenclature \u03b1 bending angle \u03c3HBS hybrid shear bonding strength d width of die P laser power R radius of punch t0 initial sheet thickness T forming temperature v scanning speed Within this work, the additive manufacturing process is conducted prior to the warm bending operation (Fig. 1). An AE with a diameter of 5 mm and a height of 5 mm is built by laser beam melting on a Ti-6Al-4V sheet metal part with a thickness of 1.0 mm or 1.5 mm. Since in prior experiments, the highest shear bonding strength for 1.5 mm sheet thickness is reached for 400 W and 300 mm/s [5], the experiments in this work are conducted with this laser parameter set. The heating of the process chamber during the LBM is set to 200 \u00b0C. In order to reduce residual stresses after the additive manufacturing process, the hybrid parts are heat treated (HT) at 850 \u00b0C for 2 hours",
            " Since Ti-6Al-4V shows low formability at room temperature [7], a special tool for bending at elevated temperature is used. Punch and die can be heated by heating cartridges. Moreover, the tool possesses the facility for heating the specimens before the bending operation by heating plates. The tool for bending of hybrid parts is described in detail in [4]. For investigating the influence of the stress state at the additively manufactured element on the shear bonding strength, the AE is placed in the tension zone (Fig. 1a) and the pressure zone (Fig. 1b) of the bending specimen. The bending operations are conducted at 400 \u00b0C with punch radii 3 mm and 7 mm and die width of 12 mm and 16 mm, to analyze the influence of the forming parameters on the mechanical properties. The bending angle for both tool combinations is 14 \u00b0 \u00b1 1\u00b0 to keep geometric deviations as small as possible. This bending angle is proven within prior experiments to be surely reached for the investigated tool combinations. In order to enable the bending of the sheet with an AE in the pressure zone, the punch has to be adapted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000413_s12206-011-1201-6-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000413_s12206-011-1201-6-Figure1-1.png",
        "caption": "Fig. 1. Configuration of the air blower.",
        "texts": [
            " On such basis, new TMM and vibration equations of the bearing-rotor system taking into consideration rolling element bearing structures were established. An actual air blower was analyzed with the new TMM to show that the bearing structure significantly affects the vibration characteristics of a bearing-rotor system. The rotor was also analyzed by FEM and results were verified, which shows that the new method proposed for vibration characteristics calculation of air blowers in this paper is credible. Fig. 1 shows an air blower used in a power station in China. It is mainly composed of a motor, pulley and bearing-rotor. The most important part of the blower is its rolling bearingrotor system. During the vibration analysis of bearing-rotor system, the foundation is treated as rigid and its vibration is ignored. Fig. 2 shows the geometry parameters of the rolling bearing-rotor system, which is composed of an impeller, pulley, shaft and a pair of tapered roller bearings (30218). The impeller, located in the middle of the two bearings, is composed of several bar pieces of weight 264"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002577_fpmc2015-9540-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002577_fpmc2015-9540-Figure1-1.png",
        "caption": "Figure 1- (a) Exploded view of an EGM. (b) Principle of operation of an EGM",
        "texts": [
            " INTRODUCTION External gear machines (EGMs) are one of the most used positive displacement machines available on the market and their success is fostered by several qualities such as compactness and low cost, a relatively high efficiency and remarkable reliability. In traditional fluid power applications, EGMs find application in both mobile and fixed applications, such as (but not only) in agricultural, construction machines and hydraulic presses. EGMs are also used in non-traditional fluid power fields, such as for fuel injection, washing or engine lubrication systems. The principle of operation of EGMs (Fig. 1) is quite simple: the two gears contained in the casing are responsible for the transport of the fluid from the inlet port to the outlet, and the meshing action between the gear teeth causes the fluid to be displaced. The reduced number of components that makes EGMs cheap units and easy to manufacture, has a drawback from the design standpoint. In fact, each main component of the unit (the gears, the sliding bushing blocks) accomplish multiple functions of positive displacement machines (volume displacement due to machine kinematic, sealing of the displacement chambers, bearing the loads induced by fluid pressure)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002681_j.jmatprotec.2016.05.029-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002681_j.jmatprotec.2016.05.029-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of specimen for X-ray fluoroscopic observation during laser cutting",
        "texts": [
            " In this study, we aimed to perform fluoroscopic observation of laser cutting fronts in real time by means of an X-ray transmission system. For this, we utilized mild steel, a widely used industrial material, together with oxygen assist gas. Also, the effect of cutting parameters (cutting speed, etc.) on cutting front profile and melt ejection behavior are discussed. This study examined the laser cutting of a mild steel (JIS SGD3M rolled carbon steel for cold-finished steel bars) with an oxygen assist gas. The chemical composition of this steel is shown in Table 1. Specimen dimensions were 4 \u00d7 120 \u00d7 4 mm (Fig. 1), set in consideration of the limited penetrative power of the X-rays. Two holes of 2.2 mm in diameter were drilled, one on each end of the specimen, for fastening to a jig. Another hole, this one also of 2 mm diameter, was drilled (not pierced) to provide a starting point for the laser cut. After surface degreasing with ethanol, the specimen was cut by laser in the longitudinal direction. A schematic diagram of the experimental setup is shown in Fig. 2. Laser cutting equipment was used together with fluoroscopic observation equipment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000056_j.precisioneng.2014.11.008-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000056_j.precisioneng.2014.11.008-Figure5-1.png",
        "caption": "Fig. 5. (a) The contacting face of the spac",
        "texts": [
            " The data for the bearings used in this study, including required bearing preload and limiting rotational speed based on preload level, were adopted from the manufacturer\u2019s catalogue, assisted in the Table 1 [18]. The axial stiffness of bearings depends directly, but nonlinearly, on the axial preload. The bearings were simplified as elastic foundation on the contact face of the spacer and bearing, as er with bearing, (b) The back face of the spring. s c i t t r t s m t t f p f c i 7 c c o b a f T R o T T hown in Fig. 5a, with an average constant axial stiffness. Frictional ontact between the spacer and spring was taken into account by nserting contact elements on the surface. Since the spring and spacer are mounted on the spindle at tight olerances, they can only move in axial direction. Therefore in he FE model, non-axial degrees of freedom of components were estrained using appropriate displacement boundary conditions. The finite element mesh was built using 3D brick elements, and he size of elements was chosen based on convergence study. Fig. 6 hows the finite element mesh used for this analysis. Once the FE odel was built, the external load on the system was applied in wo steps. First, an initial preload force was induced in the sysem by applying displacement boundary conditions on the back ace of the spring (see Fig. 5b), representing tightening of the claming nut. Then, the assembly was rotated at angular speeds ranging rom 1000 rpm up to 7000 rpm in steps to investigate the effects of entrifugal force on preload variation. The objective of the design was set to reduce the preload from ts high initial value of around 1500 N to a moderate value, around 00 N, at 7000 rpm. In order to achieve this goal, the amount of entrifugal masses was changed systematically to meet this design riterion. Fig. 7 and Table 2 report the results of the FE analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001146_0954406217722380-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001146_0954406217722380-Figure2-1.png",
        "caption": "Figure 2. Transducers arrangement: (a) transducers for cage; (b) photo of transducers for shaft.",
        "texts": [
            "5mm depth and do not reach the outer raceway in circumference to make the eddy transducer probes closer to the cage surface. The cage is also specially designed to expose larger surface area to ensure the probes to transduce the perfect cage motion signals. The radial motions of the cage are measured by two eddy current probes (yc, zc) installed in the bearing house 90 circumferentially apart and in the corresponding grooves in the outer ring. The axial motions are also measured by two probes (xc1, xc2) fixed on a panel and parallel to the bearing axis, 180 apart, as shown in Figure 2(a). The displacements of the shaft in the horizontal and vertical directions are measured by two probes (yr, zr) nearby the disc as shown in Figure 2(b). Dynamic models of rotor system and bearing cage The shaft vibrations excited by rotor unbalance can bring out the time-varying reaction forces on the supports, which can be eventually transmitted from the inner race to other bearing components including the cage. In order to predict the cage motions due to rotor unbalance theoretically, the dynamic models of the rotor system and the bearing cage are presented. Dynamic equations of the rotor system The rotor system for the scaled test rig has two support assemblies including angular contact ball bearing, bearing house, squirrel cage, as shown in Figure 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002485_s00170-016-8430-x-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002485_s00170-016-8430-x-Figure3-1.png",
        "caption": "Fig. 3 Inside blade modelFig. 2 Hypoid gear teeth cross-section",
        "texts": [
            " For a given design of a hypoid gear teeth, curvature along the lengthwise direction \u03bag,k and the flank angles \u03b1g,k at the projections of the mean point Mk on the gear teeth are given. Average radius of the cutter Rc can be chosen from the AGMA standards [19] using the following relation, cos\u03b1g;i \u03bag;i < Rc < cos\u03b1g;o \u03bag;o \u00f05\u00de Xie blade model [17] is represented in a coordinate system Ob,k, for the inside or outside blade. A unit vector sce ! k is considered along the side cutting edge, which can be represented as the function of back rake angle \u03b1o,k, side rake angle \u03b1f,k, blade angle \u03b1b,k and side relief angle \u03b3o,k, as shown in the Fig. 3. Mathematically, sce ! k is given by sce ! k \u00bc f \u03b1o;k ;\u03b1 f ;k ;\u03b1b;k ; \u03b3o;k : \u00f06\u00de Generalized parametric equation of a side cutting edges of the inside and outside blades is given by SCE ! k uk\u00f0 \u00de \u00bc uk : sce ! k \u00f07\u00de where uk is the parameter of side cutting edges. Consider a coordinate systems Ob,i and another coordinate system Ob,o parallel to the each other, in such a way that Ob,o lies along the positive Zb,i axis, at a distance equal to the point width Pw of the gear tooth, as shown in the Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000961_s10237-017-0874-x-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000961_s10237-017-0874-x-Figure12-1.png",
        "caption": "Fig. 12 Stress contour (Pa) for the central slip in the flexion movement at the PIP joint: a fifth-order Ogden, b invariants model",
        "texts": [
            " 6a and 7a and Table 1 among the Ogden models, the fifth order had the best-fitting results for the simple tension test. Thus, in order to consider the constitutive model effect, in addition to the invariantsmodel the fifth-orderOgdenmodels were also used in the FE analysis. The invariants model was implemented inABAQUS/Explicit using aVUMATuser subroutine. Also, for both the central and terminal slips, the approximate value of K was 1E+6. The results of the simulations for the central slip part in the flexion movement at the PIP joint are presented in Fig. 12. As the fifth-order Ogden model exhibited the less accurate predictions in the simple tension test compared to the invariants model and the higher values in the prediction of the pure shear (Figs. 6b, 11b) and balance biaxial tension (Figs. 6c, 11c), the von Mises stress in the Ogden model (Fig. 12a) is higher than the invariants model (Fig. 12b) under the same flexion. Also, the locations of the largest stress in both models showed different patterns, because as Fig. 6b, c indicates, the pure shear and balance biaxial tension modes of the Ogden model show the asymptotic behavior at lower stretches. Consequently, in stress analysis of the tissue under the flexion movement, this behavior may lead to miscalculations in prediction of stress distribution under flexion. The results of the simulation of the terminal slip part in the flexion movement at the DIP joint are presented in Fig. 13. As discussed in Sect. 3.2, the fifth-order Ogden model represents less stresses in pure shear (Figs. 6b, 7b) and balance biaxial tension (Figs. 6b, 7c) modes for terminal slip than central slip. As a result, the von Mises stress in Fig. 13a is less than Fig. 12a. According to Fig. 13c, the dominant mode of deformation is shear in the yz plane because the terminal slip behaves like a short beam. The deformation measure of Fig. 7 (horizontal axis) is stretch. But, due to shear modes in Fig. 13c, the principal values of strain tensor should be computed. Therefore, Fig. 13d shows the contour of maximum principal nominal strain. Converting the principal nominal strain to stretch using \u03bb = 1 + \u03b5, from Fig. 13d, it can be concluded the maximum stretch is 1.08906"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002485_s00170-016-8430-x-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002485_s00170-016-8430-x-Figure7-1.png",
        "caption": "Fig. 7 Equivalent cutting edge angle (ECE angles)",
        "texts": [
            " c k : \u00f011\u00de Since the projection pointsMk are at the same height as that ofM as shown in Fig. 6, therefore, the parameters of the curve vM,k at the height hM and then coordinates of the projection point Mk are determined in coordinate systemOb, which are given by Mb k \u00bc P !b k vM ;k \u00bc \u2212xbk vM ;k 0 zbk vM ;k 1 2 664 3 775 \u00bc \u2212hM 0 zbk vM ;k 1 2 664 3 775 : \u00f012\u00de Using the unit tangents t !b k at the mean point projections, slopes of the ECE curves are determined using the angles \u03b1e,k, are called ECE angles and are shown in Fig. 7. The unit tangents t !b k at mean point projections Mk are given by t !b k vM ;k \u00bc d P !b k vk\u00f0 \u00de . dvk d P !b k vk\u00f0 \u00de . dvk vk\u00bcvM ;k : \u00f013\u00de The ECE angles are measured between unit tangents t !b k and the unit vector x!b along the Xb axis, which are given by \u03b1e;k \u00bc cos\u22121 t !b k vMk : x!b : \u00f014\u00de There is a symmetric axis constructed for ECE tangents and its angles, angle between symmetric axis s!e and x!b is given by \u0394\u03b1e \u00bc \u03b1e;o\u2212\u03b1e;i =2: \u00f015\u00de As mentioned earlier in this article, during the face-milling process, the cutter sweep surface and the gear teeth surface remains tangent to each other at the mean point projections"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003876_s11340-019-00506-2-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003876_s11340-019-00506-2-Figure1-1.png",
        "caption": "Fig. 1 (a) Samples that were loaded along the Z-axis (i.e., the build direction) were made by removing disks of a desired height from the as-received cylinder. Samples loaded perpendicular and at a 45\u00b0 to the Z-axis were made by cutting the original rod seen in (b) and (c), respectively",
        "texts": [
            " Rods of Ti-6Al-4V with a nominal diameter of 8 mm were made by direct metal laser melting (DMLM), and obtained from GPI Prototype & Manufacturing Services (Lake Bluff, IL). DMLM is a powder bed fusion process that involves fully melting the powder to create a highly dense part (in contrast to direct metal laser sintering, DMLS, which generally achieves lower temperatures). Following the terminology defined in ISO/ASTM 52921 [36], the build direction is denoted as the Z-direction, as illustrated in (Fig. 1(a)). The X- and Y-axes, which align with the edges of the DMLM machine, were not provided by the manufacturer, so will be arbitrarily chosen here. For the material used in this study, the thickness of each build layer is 30 \u03bcm. Different cylindrical (disk) samples for use in quasi-static and dynamic compression experiments were made by electrical discharge machining (EDM) of the as-received cylinder such that load could be applied either along the Z-axis (i.e., the build direction), perpendicular to the Z-axis, or at an angle of 45\u00b0 to the Z-axis",
            " The samples that were loaded along the Z-axis were simply cut by removing the desired length from the initial cylinder. For the quasistatic compression experiments along the Z-axis the cylindrical samples were of diameter 8 mm and length ranging from 5 to 8 mm, while for the dynamic experiments they were disks of diameter 4 mm and thickness 2 mm. The remaining two loading-direction geometries were more complicated to extract from the initial rod in that the EDM could not make a cut parallel to the face of the initial cylinder. Instead, a piece of Ti-6Al-4V was removed by EDM as illustrated in (Fig. 1b, c) and that piece was further machined to have parallel faces. Finally, some of the samples where the load was to be applied along the build direction were annealed at 850 \u00b0C for five hours and then air cooled, similar to the annealing procedure used in [37]. The complex thermal history of the additive manufacturing process, in which material is repeatedly heated and cooled, results in residual stresses throughout the manufactured part. The purpose of annealing was to be able to compare the effects of reducing any residual stresses that may have been otherwise present in the AMmaterial"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001928_icit.2015.7125147-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001928_icit.2015.7125147-Figure5-1.png",
        "caption": "Fig. 5. Concept of modulated potential field.",
        "texts": [
            " Epipolar geometry. Fig. 4. Stereo system with parallel optical axes. yl = yr (6) Fig. 4 shows triangulate geometry in stereo system with parallel optical axes. Corresponding points in stereo images can be reconstructed to three dimensional space using (7) to (9). x = xl + xr 2 L xl \u2212 xr (7) y = yl L xl \u2212 xr (8) z = f L xl \u2212 xr (9) In this paper, modulated potential field for position adjusting with human interaction is defined. The required performances of modulated potential field are shown in Fig. 5. There are three requirements that described as follow. 1) Far from the Destination: Surgeon can freely operate the manipulator. And the trajectory will be decided by human. 2) Approaching to the Destination: The manipulator will apply adjusting force for correctly converging to the destination. 3) Passing the Destination: The manipulator will apply adjusting force for collision avoidance. The Lennard-Jones potential (L-J potential) field can be suitable with these three requirements. It is a mathematically simple model that approximates the interaction between a pair of neutral atoms or molecules"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000599_j.engfailanal.2012.02.008-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000599_j.engfailanal.2012.02.008-Figure7-1.png",
        "caption": "Fig. 7. Computational model of casting pedestal traverse.",
        "texts": [
            " Subsequently, it was found out that one bolt was destroyed and in bodies of several bolts were identified cracks. These failures were repaired \u2013 cracks were abraded and welded (Fig. 6), damaged bolts were replaced by new ones (bolts M48 were replaced by bolts M52). At the same time was made decision that it is necessary to accomplish stress analysis of traverse of casting pedestal. Computational model of traverse (due to symmetry only one half of beam and middle part with arm of toothing was considered) is given in Fig. 7. Computation of traverse was accomplished for all load cases given above. In Fig. 8 is given field of equivalent stresses on traverse of casting pedestal for the most danger state of loading. In Figs. 9 and 10 are shown details of equivalent stresses on upper side of traverse beam. For experimental determination of time-dependent changes of stress in the traverse beams was used the electrical resistance strain-gage method [3,6,8]. Strain-gage measurements were realized by strain-gage apparatus SPIDER with application of strain-gages produced by company HBM according to Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002931_aa591b-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002931_aa591b-Figure11-1.png",
        "caption": "Figure 11. The bearing tested rig.",
        "texts": [
            " In figure\u00a0 9(b), the impulsive features of 32 Hz J13and its two harmonics of 64 Hz and 96 Hz are identified. In figure\u00a010(b), the impulsive feature of 32 Hz and its harmonics e.g. 64 Hz, 96 Hz and 128 Hz, are extracted. In comparison with the CMM method, the results show that the OMM method has superior performance in extracting the impulsive features. In order to verify the effectiveness of the proposed method, a benchmark bearing dataset, which was provided by the Case Western Reserve University (CWRU) [28] Bearing Data Center, is adopted. As shown in figure\u00a0 11, the test stand is made up of a 2 HP motor, a torque transducer, a dynamometer, and the related control electronics. The tested bearing SKF 6205, whose geometry parameters are given in table\u00a01, supports the motor shaft. Single point faults on the inner and outer races were introduced into the experiment using electrodischarge machining. Vibration data was collected using accelerometers, which were placed at the 12 o\u2019clock position at the drive and fan ends of the motor housing and at the motor supporting base plate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003962_978-3-030-20131-9_196-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003962_978-3-030-20131-9_196-Figure1-1.png",
        "caption": "Fig. 1. Robot design: a) Planar mechanism. b) End effector. c) Detail of the transmission.",
        "texts": [
            " The mechanism proposed here has a 5R topology where the two bars joined at the end effector are the ones that are flexible, made of Nylon 6 with a length of 790 mm and a section of 10 mm of diameter. The manipulator is thus regarded as collaborative, as the flexibility of the bars should reduce the damage to the human in situations of impact and entrapment, although no experimental tests have been done in this aspect. These bars are fixed at 115\u00ba to the two rigid bars connected to the rotary actuators, made of carbon fibre. In Fig.1a, it is seen the whole design, and in Fig.1b and Fig.1c there is a detail of the end effector and the actuators and fixed support. The links between the actuators and bars, the fixed joint between the flexible and carbon fibre bars and the end effector are all made by 3D printing in ABS+. To complete the prototype definition, in Fig.2a the mechanism diagram is seen in the position where the two bars do not deflect, which has been taken as a home position. In Fig. 2b. the mechanism is in a generic position. In the Cosserat bar model, a parametric curve in the space defines the position of the center of the deformed bar p(s), and a rotation matrix R(s) defines the orientation of the transverse rigid section, where s is the arc length, see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002737_9781118899076-Figure8.19-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002737_9781118899076-Figure8.19-1.png",
        "caption": "FIGURE 8.19 Orientation of electron spins in the spin-frustrated ground state of NI (a) and the calculated geometric structure of NI (b). Reproduced with permission from Ref. [34]. \u00a9 2011 RSC Publishing.",
        "texts": [
            " Importantly, the MCD intensity at high field does not decrease in intensity with varying temperature as 1/T, the Curie law behavior predicted for an isolated S =\u00bd ground state (Fig. 8.18d). For instance, the intensity at 25,000 cm\u22121 first decreases and then increases in intensity (Fig. 8.18e), consistent with the Boltzmann population of a low-lying excited state at approximately 150 cm\u22121 above the S =\u00bd ground state and with a different MCD signal (Fig. 8.18f ). This low-lying excited state requires all three coppers of the TNC are close to equally AF coupled (~500 cm\u22121). This situation results in a spin-frustrated ground state for NI. Figure 8.19a shows that AF coupling of Cu1 with Cu2 and Cu2 with Cu3 leads to a parallel alignment of the spins on Cu3 and Cu1, although theywant to be AF coupled due to the superexchange pathway associated with a Cu1\u2013Cu3 bridging ligand. Therefore, all three coppers must be bridged by the products of complete dioxygen reduction. Additionally, the unique g < 2.0 values, which originate from a phenomenon known as antisymmetric exchange that is associated with this spin frustration [52], and the sign and nature of the temperature dependence of the pseudo-A term in the MCD spectra in (Fig. 8.18a, bottom) unambiguously lead to the structural assignment of NI as containing a \u03bc3-oxo ligand bridging all three of the coppers of the TNC and a \u03bc2-OH between the T3s (Fig. 8.19b) where the two bridging oxygen ligands originate from dioxygen reduction [53, 54]. Having defined the electronic and geometric structures of PI and NI, the mechanism of the 2nd two-electron reductive cleavage of the O\u2500O bond was explored [49]. 184 MOLECULAR PROPERTIES AND REACTION MECHANISM (a) 6000 SF-Abs RFO-LT-MCD 10 msec 1 msec 365 nm 318 nm T1 T1 7T ^ 1.8K at 27,560 cm\u20131 31,780 cm\u20131 N1 N1 4000 2000 40 20 0 \u201320 \u201340 100 50 80 60 40 20 0 0 50 100 150 Temperature (K) Temperature dependence of MCD intensityX-band EPR 25,000 cm\u20131 1000 250 150 100 70 AE (cm\u20131)0 \u201350 \u2013100 15,000 8970 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure2-1.png",
        "caption": "Figure 2: Optimized configuration of the lattice structure in the simulation model (left) and in a 3D prototype (right) [8]",
        "texts": [
            " However, a variation of the density or the kind of structure inside of one and the same part is complex or even not possible. In [8], another approach is presented, in which plane lattice structures are getting adapted to a parts surface. This is demonstrated at the example of a deployable wing. For the optimization, an initial structure is modelled with one dimensional beam elements. This structure gets optimized by the adaption of the dimension or the elimination of the single lattice elements (see figure 2). While in the preceding works, the principle build-up of the structures was predefined, Cansizoglu [9] presents an optimization algorithm, which adapts the position and orientation of the single struts. In his work, he derives an optimization function from an initial structure, which adapts several parameters like the nodes\u2019 positions or the beams\u2019 cross sections. However, the number of beams and nodes is predefined here, too. Furthermore, the parameter based optimization algorithm is computationally intensive, so that this approach will prospectively not work for large and complex structures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000616_aim.2012.6265962-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000616_aim.2012.6265962-Figure4-1.png",
        "caption": "Fig. 4. Example of jumping motion",
        "texts": [
            " That is, contacting-foot may slip forward or backward depending on the foot\u2019s velocity in traveling direction. As shown in Fig. 2, we distinguish contacting patterns by changing the dimension of state variables. That is, although we do not address the situation that supporting-foot slips or both feet are in the air, Eq. (16) can also represent these dynamics: adding position variable of walking direction y0 to q in Eq. (16) when supporting-foot begins slipping; and jumping motion by adding further variable of upright direction z0 to q when jumping represented by Fig. 4. Table II indicates all possible walking gaits regarding contacting situations\u2014surface-contacting (S), point-contacting (P) and Floating (F)\u2014of supporting-foot (S.F.) and floatingfoot or contacting-foot (F.F.). The Table is basically divided into three blocks representing the gait\u2019s varieties from a point of states of supporting-foot, such as, \u201cStop,\u201d \u201cSlip\u201d and \u201cAir.\u201d Figure 5 depicts all possible gait transition of bipedal walking based on event-driven, which indicate that appropriate dynamics and variables are selected and applied according to the phase or state, which are listed in Table II"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001165_icuas.2017.7991362-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001165_icuas.2017.7991362-Figure3-1.png",
        "caption": "Fig. 3. Forces and moments on the UAUV in an octo-quadcopter configuration.",
        "texts": [
            " We can observe that the drag forces are highly nonlinear and hard to model, but in general they act as a natural damper in the system. We select a multirotor in an octo-quadcopter configuration, as depicted in Fig. 12. It is composed of four coaxial rotor pairs, four rotors spinning clock-wise and four spinning counter clock-wise in order to cancel out the reactive torques. This kind of configuration is ideal for UAUVs to transition between mediums, since the top rotors can spin at different speeds than the bottom ones, according to the medium they are immersed in [7]. The forces and moments exerted in the vehicle (see Fig. 3) are [ T \u03c4 ] =  1 1 1 1 1 1 1 1 \u2212l l l \u2212l l \u2212l \u2212l l l l \u2212l \u2212l l l \u2212l \u2212l \u2212b b \u2212b b b \u2212b b \u2212b   f1 ... f8  (12) and fi = \u03c9 2 i \u03c1Km (13) where fi, \u03c9i are respectively the force and the angular speed produced by rotor i. The constants l, b and Km stand for the distance from the rotor to the center of mass of the vehicle, an angular speed to torque constant and the motor\u2019s constant, respectively. Once more we note the influence of the medium density in the generation of forces and torques by the motors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003076_tmag.2017.2708748-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003076_tmag.2017.2708748-Figure9-1.png",
        "caption": "Fig. 9. Temperature distribution. (a) SMTLOM. (b) SMTHLOM.",
        "texts": [
            "5 times the rated value, i.e. enough margin has been spared for safe operation. C. Thermal Distribution In a practical device, e.g. a linear compressor, the operation frequency is not very high (usually no more than 50Hz). The core loss is much smaller than the copper loss. That is, the copper loss could be approximately regarded as the main heat source of the LOMs. The thermal distribution is analyzed by using the 3-D FEA, with the dissipation mode regarded as closed natural cooling and the ambient set as 25\u2103. Fig. 9 depicts the temperature distribution of the two machines. It is shown that with the same amount of copper loss (13W for both two LOMs), the temperature distributions are almost the same. The temperatures of both NdFeB and ferrite magnets are all below the critical values, which sets relative large temperature margins to protect the magnets from irreversible demagnetization and guarantees the safe operation. D. Material Consumption and Cost Since both LOMs have similar performance, it is a sensible practice to compare their material costs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000647_iciea.2015.7334244-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000647_iciea.2015.7334244-Figure5-1.png",
        "caption": "Fig. 5 Collision detection of sphere-cuboid",
        "texts": [],
        "surrounding_texts": [
            "Keywords-collision detection; manipulator; vector relation\nI. INTRODUCTION Collision detection plays an important role in manipulator field, which includes manipulator-environment collision and manipulator-manipulator collision. It is necessary to build proper geometric models to detect collision of manipulator rapidly. Therefore, how to model and how to process model is directly related to the precision and efficiency of collision detection algorithm.\nConsidering the complexity of manipulator\u2019s shape and the diversity of environment, and in order to reduce mathematical operator to ensure efficiency, bounding box algorithm is generally used in collision detection. Basic idea of the algorithm is to use a slightly larger geometry with simple feature(called the bounding box) to represent complex objects approximately. Simple geometry such as sphere and cuboid are widely used. Angel P. del Pobil [1] used hierarchies of detail based on spheres to represent complex objects, which was of high efficiency. Fan Ouyang [2] analyzed the advantages of hierarchical structure and octree data structure, and proposed an octree-based spherical hierarchical model for collision detection. Crnekovic Mladen [3] took the environment as a series of sphere or prism and compared the efficiency and error between the two models. Kei Okada [4] adopted AABB algorithm to envelope objects as closely as possible, and demonstrated real-time collision detection. Zonggao Mu [5] turned to OBB algorithm considering that most space manipulators are composed of modular and standardized\ndevices, and developed an integrated simulation system based on OSG [6,7] to verify the algorithm. Byungjin Jung [8] used a kind of band pass filter to avoid false alarm in collision detection, and the experimental result showed that the method is not only fast but also robust. Apart from bounding box algorithm, Ruining Huang [9] adopted a hierarchical bounding box method and space decomposition method comprehensively. Minxiu Kong [10] proposed a collision detection algorithm for coordinated industrial robots based on calculating the shortest distance between two geometry.\nObviously, these three bounding boxes have their own strong points and weakness. Among them, Sphere box is the simplest with high efficiency for collision detection, although with low precision at the same time. AABB is acceptable in both efficiency and precision. OBB is of high precision, but it need to detect 15 separate axes to determine collision status in 3D environment. To ensure both precision and efficiency, in this paper a rapid collision detection algorithm for manipulator based on the vector relation of point, line segment and rectangle is proposed. Besides, we put forward the concept of \u201ccenter distance\u201d and \u201cgeometric dimension\u201d, and provided a center distance criterion to detect collision rapidly.\nII. GEOMETRY OF MANIPULATOR AND ENVIRONMENT First, we simplified manipulator and environment based on the bounding box theory. Considering that most manipulators consist of modular components, we adopted cylinder and cuboid to represent a manipulator. For diverse obstacles in environment, we used sphere, cylinder, cuboid and combination of them to represent.\n934978-1-4799-8389-6/15/$31.00 c\u00a92015 IEEE",
            "Therefore, object collision is simplified to geometry collision. In this paper, geometry is subdivided as the combination of point, line segment (denoted as line for convenience) and rectangle. Six kinds of geometry (point, line, rectangle, sphere, cylinder and cuboid) above are collectively called basic geometry, as shown in Fig. 2.\n0P\ncy\ncx\ncz\ncO\n2P\n1P 3P 4P\n5P\n6P 1s\n2s3s\n1P\n2P cO 1s\n2s\n0P\n1P\n2P r\nr\nFig. 2 Basic geometry\nFurthermore, we extracted concept of \u201ccenter distance\u201d \u2013 the distance between center of two geometry; and \u201cgeometry dimension\u201d \u2013 the maximum distance from center to surface of a geometry. The information of all basic geometry is shown in Table 1.\nIII. COLLISION DETECTION OF BASIC GEOMETRY After simplification of manipulator and environment, object collision can be classified into 4 kinds of collision \u2013 spherecylinder, sphere-cuboid, cylinder-cylinder and cuboid-cuboid collision. For convenience, cylinder-cuboid collision is seen as cuboid-cuboid collision. To skip collision detection between 2 geometry far away, we proposed a center distance criterion: if the center distance between 2 geometry is greater than the sum of their geometric dimension, there is no collision. Otherwise it need further detection. In addition, considering safety margin dS , the criterion is shown with following expression:\n( , ) . . i j i j Sif Dist G G D D d No collision else Need more detection > + +\nIn further detection, the 4 types of collision detection can be converted into 5 basic detections: point-line, point-cuboid, lineline, line-rectangle, line-cuboid. Their relation is given in Fig. 3.\nThe following is an introduction of the 4 types of collision detection. The algorithm returns 1 if any collision happens,; otherwise returns 0.\na) Apply center distance criterion on sphere-cylinder. The algorithm returns 0 if (1) meets, otherwise turns to b);\n2 2 0 1 2( ) = || || || || /4S C m C S SDist G ,G P P PP r r d> + + + (1)\nb) Compute the shortest distance dmin between point P0 to line P1P2.\nDo projection of point P0 to line P1P2. If 1 1 2/ [0,1]vPP PP\u03bb = \u2208 , then min 0|| ||vd P P= , otherwise\nmin 0 1 0 2 min{ || ||,|| ||}d P P P P= . c) Detect collsion of sphere-cylinder by dmin. The algorithm returns 0 if (2) meets; otherwise returns 1.\na) Apply center distance criterion on sphere-cuboid. The algorithm returns 0 if (3) meets, otherwise turns to b);\n2 2 2 0 1 2 3Dist( , ) || || || || || || || ||S C c S SG G P O r d= > + + + +s s s (3)\nb) Detect whether point P0 is outside of cuboid C. For any surface of cuboid with outward normal vector in\nand center point iP , if 0 0i iP P n\u22c5 < , then point P0 is outside, otherwise turns to c). The algorithm computes dmin of point P0 to the corresponding surface. The algorithm returns 0 if dmin > rS ; otherwise returns 1.\nc) Detect next surface.\n2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA) 935",
            "If all the 6 surfaces of cuboid fail to meet 0 0i iP P n\u22c5 < , then point P0 is inside of cuboid, which means 1 to return. 3) Cylinder \u2013 Cylinder\n1P\n2P\n1vP\nmind\n3P\n4P\n2vP\n1C\n2C\n2l 1l\n1mP 2mP\n1Cr 2Cr\n1mP\n4P\n2mP\n1vP\n2vP\nmind\n2P\n1P\n1 3 1 4 2 3 2 4 min{|| ||,|| ||,|| ||,|| ||}PP PP P P P P . c) Detect collsion of cylinder-cylinder by dmin. The algorithm returns 0 if (7) meets; otherwise returns 1.\na) Apply center distance criterion on cuboid-cuboid: The algorithm returns 0 if (8) meets, otherwise turns to b);\n2 2 2 1 2 1 2 3\n2 2 2 1 2 3\nDist( , ) || || || || || || || ||\n|| || || || || ||\nC C c c\nS\nG G O O '\n' ' ' d\n= > + +\n+ + + +\ns s s\ns s s (8)\nb) Detect whether vertex of C1 is inside of C2. The collision is the point-cuboid type. The algorithm\nreturns 1 if a vertex collides, otherwise detects next vertex. If all 8 vertexes have no collision, turns to c);\nc) Detect whether edge of C1 is in collision with C2. The collision is the line-cuboid type. The algorithm\nreturns 1 if an edge collides, otherwise detects next edge. If all 12 edges have no collision, returns 0.\nIV. COLLSION DETECION OF MANIPULATOR In the base of simplified geometric models and collision detection of basic geometry, we realized collision detection of manipulator. The algorithm flow chart is shown in Fig. 8.\nThe simplified model of a n-DOF manipulator is shown in Fig. 9.\nTo process collision detection of manipulator, we need the key information of basic geometry \u2013 key point, key vector and key length in an absolute reference system. In this paper, we took the base coordinate system of manipulator as the absolute reference system. Different geometry demands for different key information, as listed in Table 2. The expression of key point and key vector are same in Cartesian form. When joint angles are given, the relevant configuration, key point and key vector of manipulator can be solved by forward kinematics. The key length will not change without considering flexibility of manipulator.\n936 2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA)"
        ]
    },
    {
        "image_filename": "designv11_13_0003621_j.apm.2018.10.002-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003621_j.apm.2018.10.002-Figure1-1.png",
        "caption": "Figure 1: General sketch of the new screw-based sub-problem 1",
        "texts": [
            " In the following sections, IK models of the newly defined sub-problems are established according to basic geometrical descriptions, where two simple evidences would be used. First, when a point rotates about an axis, its distance to any point on the rotational axis remains unchanged; and second, vectors located on the motion trajectory plane are perpendicular to the axis of rotation. Based on the screw-based descriptions given in Section 2.1, IK models of three new sub-problems are presented in the following parts. The general sketch of the new screw-based sub-problem 1 is shown in Figure 1, where \u03b8i, qi are the angle of rotation about axis i and one point on this axis, respectively; a and b denote the initial and target position, respectively; cN1 and cN2 are the process points; and the dotted circles are the movement trajectories. The FK for each case can be described as follows: exp (\u03be\u0302N1\u03b8N1) exp (\u03be\u0302N2\u03b8N2) exp (\u03be\u0302N3\u03b8N3)a \u2032 = b \u2032 (6) The inverse motion from point b to point a can be decomposed into three rotational motions: rotation about joint N1 from point b to point cN2, rotation about joint N2 from point cN2 to point cN1, and rotation about joint N3 from point cN1 to point a",
            " Based on two simple geometrical evidences, the following equation set can be obtained for the new screw-based sub-problem 1:    \u2016qN1 \u2212 b\u2016 = \u2016qN1 \u2212 cN2\u2016, (b\u2212 cN2)\u03c9T N1 = 0 \u2016qN2 \u2212 cN2\u2016 = \u2016qN2 \u2212 cN1\u2016, (cN2 \u2212 cN1)\u03c9T N2 = 0 \u2016qN3 \u2212 cN1\u2016 = \u2016qN3 \u2212 a\u2016, (cN1 \u2212 a)\u03c9T N3 = 0 (7) where the symbol \u201d\u2016 \u00b7 \u2016\u201d denotes the calculation of the length of one vector, and \u03c9N1, \u03c9N2, and \u03c9N3 are unit rotational vectors in the axial directions of joints N1, N2, and N3 in the robotic base frame, respectively. After finding cN1 and cN2, the problem can be converted into three PK-1 sub-problems, which can be expressed as:    exp(\u03be\u0302N1\u03b8N1)c \u2032 N2 = b \u2032 exp(\u03be\u0302N2\u03b8N2)c \u2032 N1 = c \u2032 N2 exp(\u03be\u0302N3\u03b8N3)a \u2032 = c \u2032 N1 (8) Therefore, the three angular displacements in Figure 1 can be solved based on the solution to the PK-1 sub-problem described in literature report [24]: \u03b8m = arctan 2(\u03c9m(u\u2217 m \u00d7 v\u2217m)T ,u\u2217 mv \u2217 m T ),m = N1, N2, N3 (9) ACCEPTED M The general sketch of the new screw-based sub-problem 2 is presented in Figure 2. This time, the inverse motion can be decomposed into four rotational motions: rotation about joint N1 from b to cN3, rotation about joint N2 from cN3 to cN2, rotation about joint N3 from cN2 to cN1, and rotation about joint N4 from cN1 to a. The forward description of the inverse motion from b to a can be given as: { exp(\u03be\u0302N1\u03b8N1)RN2 exp(\u03be\u0302N3\u03b8N3) exp(\u03be\u0302N4\u03b8N4)a \u2032 = b \u2032 , if RN2 is given exp(\u03be\u0302N1\u03b8N1) exp(\u03be\u0302N2\u03b8N2)RN3 exp(\u03be\u0302N4\u03b8N4)a \u2032 = b \u2032 , if RN3 is given (10) where RN2 and RN3 present the rotational motions about axes N2 and N3, respectively, and for the robotic IK problems they are equal to a single joint motion or the sum of two or more joint motions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002547_1.4033101-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002547_1.4033101-Figure4-1.png",
        "caption": "Fig. 4 Outer ring cross section with vectors locating outer raceway centroids in body fixed reference frame",
        "texts": [
            " The inner rings are split in the middle to allow assembly. The use of a common outer ring requires modification of the contact routine between a ball and outer raceway. The centroid of the outer raceway is not located at the center of mass of the outer ring. An additional vector is needed to locate the outer raceway centroid with respect to the outer ring center of mass. The position of each outer raceway centroid is defined with respect to the center of mass of the cartridge outer ring by vectors in the body fixed reference frame as shown in Fig. 4 012201-2 / Vol. 139, JANUARY 2017 Transactions of the ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jotre9/935372/ on 02/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use r bf RR1 \u00bc e; 0; 0\u00bd T (1) r bf RR2 \u00bc \u00fee; 0; 0\u00bd T (2) The position vector subscripts refer to the ring and raceways 1 and 2 and the superscript refers to the body fixed frame. The value e is the offset distance of the outer raceways from the outer ring center of mass"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001568_j.mechmachtheory.2018.07.009-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001568_j.mechmachtheory.2018.07.009-Figure3-1.png",
        "caption": "Fig. 3. Schematic of the planar five-bar manipulator.",
        "texts": [
            " (42) states that at a gain-type singularity, the active joint velocity lies on the tangent space to the gain-type singularity manifold in the actuator-space . This is a necessary condition for any motion that encounters gain-type singularity, whether it crosses the same (in the task-space), or stays on it. The example of the planar five-bar manipulator (described in detail in Section 4.1.1 ) allows an easy visualisation of the condition. When the link lengths are chosen to be l 0 = 1 , l 1 = 2 / 3 , l 2 = r 1 = r 2 = 1 / 3 , (see Fig. 3 ), the five-bar manipulator becomes an equivalent four-bar mechanism with a coupler length of r 1 + r 2 at a gain-type singularity. It can continue on the singularity manifold, until, eventually, it comes to the shape of a triangle, i.e., a pose where the equivalent four-bar itself becomes singular. The motion is animated (kinematically) in a video clip (filename: fivebar1.mp4 ) associated with the paper, and the same is depicted in the \u03b81 \u2212\u03b82 space in Fig. 1 . It can be seen that the path meets the singularity manifold (i",
            " On the other hand, the latter, being a six-DoF spatial manipulator, demonstrates the generic nature of the formulations. The geometric features and kinematics of the manipulators, derivation of various terms in the equation of motion, the details of the trajectories tracked, and finally the measures used for validating the simulations are explained in this section. The geometry, kinematics and mass properties of the five-bar manipulator are described below in brief. The planar five-bar manipulator is shown schematically in Fig. 3 . In this case, \u03b8 = [ \u03b81 , \u03b82 ] , and \u03c6 = [ \u03c61 , \u03c62 ] . The loop- closure equations are of the form \u03b7 = [ \u03b71 , \u03b72 ] = [0 , 0] , where: \u03b71 = l 1 cos \u03b81 + r 1 cos \u03c61 \u2212 l 0 \u2212 l 2 cos \u03b82 \u2212 r 2 cos \u03c62 , (53) \u03b72 = l 1 sin \u03b81 + r 1 sin \u03c61 \u2212 l 2 sin \u03b82 \u2212 r 2 sin \u03c62 . (54) The matrices J \u03b7\u03b8 and J \u03b7\u03c6 are given by: J \u03b7\u03b8 = [ \u2212l 1 sin \u03b81 l 2 sin \u03b82 ] , J \u03b7\u03c6 = [ \u2212r 1 sin \u03c61 r 2 sin \u03c62 ] . (55) l 1 cos \u03b81 \u2212l 2 cos \u03b82 r 1 cos \u03c61 \u2212r 2 cos \u03c62 The links of the five-bar manipulator are taken to be solid cylinders of diameter 20 mm. Therefore, the mass centres of the links are at their respective geometric centres, and their inerti\u00e6 are computed accordingly. The material is assumed to be aluminium, with a density of 2700 kg/m 3 . The manipulator is assumed to be in the vertical plane. Different sets of link lengths have been used in different simulations, and hence the link lengths have been noted in the appropriate places. Additionally, a concentrated payload of m = 2 kg is assumed to be at the end-effector p (see Fig. 3 ) in all the simulations. The derivation of the equation of motion is fairly straightforward, and has hence been omitted for the sake of brevity. The SPM chosen in this work is one with two irregular hexagons serving as its fixed base and moving platforms, respectively. The dimensions of these platforms and other physical parameters, such as the mass and inertia properties of the manipulator, are listed in Appendix B . The derivation of the kinematic constraints as well as the equation of motion are described in the following sections",
            " 172 0 . 431 0 . 431 \u23a4 \u23a5 \u23a6 \u23a1 \u23a2 \u23a3 h 1 h 2 h 3 h 4 \u23a4 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a3 \u03c4\u03b81 \u03c4\u03b82 0 0 \u23a4 \u23a5 \u23a6 . (84) It is noted that rows of the sub-matrix E \u03c6 are equal (linearly dependent), but the corresponding values in Q are not equal. This clearly shows that the last two equations are inconsistent, and there exists no h that can satisfy both these equations simultaneously. This implies that the dynamic equations of motion are inconsistent for the stipulated trajectory when the actuators are present at the base pivots (see Fig. 3 ). In summary, the actuator torques cannot be computed through this method either. \u2022 Method C . Upon substitution of values for the position, velocity and acceleration coordinates in Eq. (35) , one obtains: \u23a1 \u23a2 \u23a3 0 . 586 0 . 586 \u22120 . 107 0 . 576 0 . 447 0 . 447 0 . 355 0 . 003 \u22120 . 041 \u22120 . 041 0 . 691 \u22120 . 723 \u22120 . 041 \u22120 . 041 \u22120 . 309 0 . 277 \u23a4 \u23a5 \u23a6 \u23a1 \u23a2 \u23a3 \u03c4\u03b81 \u03c4\u03b82 0 0 \u23a4 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a3 \u221226 . 243 13 . 506 54 . 353 \u221222 . 509 \u23a4 \u23a5 \u23a6 . (85) The row-reduction operation is performed on the system of equations as illustrated in the non-singular case to arrive at: [ 0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003488_j.ifacol.2018.05.085-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003488_j.ifacol.2018.05.085-Figure1-1.png",
        "caption": "Fig. 1. Helicopter free body diagram during hover and the Body fixed frame",
        "texts": [
            " The inertial frame {OI , iI , jI , kI} is the commonly used North East Down frame and the body frame {OB , iB , jB , kB} is having its origin fixed at the centre of gravity, iB towards the body head, jB towards the right of the body and direction of kB is defined by the right-handed rule. The translational velocities VB = [u v w]T and the rotational velocities \u03c9B = [p q r]T are used to describe helicopter motion with respect to the body frame. The body frame and the velocities with respect to this frame are illustrated in Fig 1. Any vector in the body frame can be mapped to inertial frame using rotation matrix which is parameterised with respect to the Euler angles roll (\u03c6), pitch (\u03b8) and yaw(\u03c8). The force and torque components acting on the helicopter in body frame are represented by fB = [X Y Z]T , \u03c4B = [LM N ]T respectively. The equations of motion of a helicopter in the inertial frame are given by the fundamental principles based on conservation of linear and angular momentum. The translational dynamic model can be summarized as f = m dV dt |I (1) where m represent the mass and the suffix I represent the inertial frame",
            " FORCES AND MOMENTS DURING HOVERING FLIGHT It is necessary to analyse hover flight regime besides its simplicity regarding the aerodynamic forces acting. It is one of the toughest flight condition due to the action of gusty air produced by the main rotor on the fuselage. The coupling between the degrees of freedom and between the control inputs makes the scenario more complicated. It requires continuous adjustments in the control inputs to remain stationary. The thrust and torque generated by main rotor and tail rotor, flapping angles of the main rotor are shown in Fig.1. The tail rotor is assumed to have force component only in y- direction. Since the vehicle is immobile with respect to the ground, forces acting on vertical/horizontal stabilizers and other 5th International Conference on Advances in Control and Optimization of Dynamical Systems February 18-22, 2018. Hyderabad, India 537 506 Femi Thomas et al. / IFAC PapersOnLine 51-1 (2018) 504\u2013511 parts can be neglected. The main and tail rotor contribution is dominating in the force and torque developed on the body",
            " While equating the force in x-direction, X = \u2212Tm sin (a) (30) Partial derivative of the force component shown in (30) with respect to linear velocity u is given by, Xu = \u2212 ( \u2202Tm \u2202u sin (a) + Tm cos(a) \u2202a \u2202u ) (31) Similarly, expressions for other stability and control derivatives are found by considering the respective force and moment equations. The main and tail rotor thrust, torque equations are obtained from the basic aerodynamics explained in Budiyono et al. (2009) and Gavrilets (2015). 5th International Conference on Advances in Control and Optimization of Dynamical Systems February 18-22, 2018. Hyderabad, India Femi Thomas et al. / IFAC PapersOnLine 51-1 (2018) 504\u2013511 507 Fig. 1. Helicopter free body diagram during hover and the Body fixed frame parts can be neglected. The main and tail rotor contribution is dominating in the force and torque developed on the body. The forces and torques experienced by the helicopter during hover condition are, fB = [ XM YM + YT ZM ] +RT [ 0 0 mg ] (22) \u03c4B = [ LM MM +MT NM ] + [ YMhM + ZMyM + YThT \u2212XMhM + ZM lM \u2212YM lM \u2212 YT lT ] (23) where hM ,yM ,lM are the main rotor coordinates and hT ,lT is the tail rotor coordinates. Force and moments with respect to the body frame in terms of the thrust generated and the flapping angles of the rotor is given by the expressions, XM = \u2212TM sin(a), YM = TM sin(b), ZM = \u2212TM cos(a) cos(b), YT = \u2212TT , LM = cbb \u2212 QM sin (a), MM = caa+QM sin (b), NM = \u2212QM cos (a) cos (b), which are obtained by the simple decomposition of forces along the references considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003153_iemdc.2017.8002352-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003153_iemdc.2017.8002352-Figure4-1.png",
        "caption": "Fig. 4. Cross-section view from the radial direction of comparative analysis models of bearingless motors.",
        "texts": [
            " To suppress decrease of torque and suspension force caused by adopting Nd bonded PM instead of Nd sintered PM can be realized by shortening magnetic gap length which will cause an explosion of eddy current loss when Nd sintered PM is employed. To confirm effectiveness of the proposed BelM, this paper performs comparison of analysis results of six BelM models using distributed winding structure, concentrated winding structure, Nd sintered PM, and Nd bonded PM, respectively. Theses BelM models have the same specification indicated in Table II and the same current density. Fig. 3 and Fig. 4 show the structure of the six BelM models. BelM models with Nd sintered PM have wide magnetic gap length of 7 mm in order to reduce eddy current loss in PM. As mentioned previously, Nd bonded PM has extremely low eddy current loss due to low conductivity. Therefore, the proposed 6- 8 pole, 9 slot concentrated winding model with Nd bonded PM has short magnetic gap length of 6 mm. It is noted that \u201c6-8 pole\u201d means six pole motor winding and eight pole suspension winding. 6-8 pole BelMs require high drive frequency of 3 kHz at rated rotational speed of 60,000 r/min",
            " It is necessary to drive the BelMs under high inverter switching frequency in experiment in order reduce the switching ripple current which cause iron loss increase. It is possible to drive a 6-8 pole BelM under high speed rotation of 60,000 r/min by using the PWM inverter equipped with next generation high performance device SiC-MOSFET. It is known that the distributed winding structure requires longer coil end space than the concentrated winding structure. A length of the coil ends can be reduced by increasing of the number of pole and slot. Each thickness of stator and coil end of the six BelM models is shown in Fig.4. In order to actualize stable magnetic levitation under high speed rotation of 60,000 r/min, high-response and high-accuracy current regulation is required. Furthermore, it is important to decrease suspension inverter capacity to reduce the product cost. Therefore, differential winding configuration, which is composed of motor winding and suspension winding, is adopted in this paper. The ratio between the motor winding and the suspension winding in the same stator slot area is decided on the condition that suspension force per one unit becomes equal at 70 N or more shown in Table I in order to achieve levitation rotation stably"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000929_b978-0-444-52215-3.00008-8-Figure8.4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000929_b978-0-444-52215-3.00008-8-Figure8.4-1.png",
        "caption": "FIGURE 8.4 Multiple-pass heat exchangers. A one-shell pass, two-tube pass (1\u20132) heat exchanger with baffles to direct the shell-side flow. The two-tube pass design reduces the number of tubes in a heat exchanger from two tubes to one tube and increases the tube-side fluid velocity so that it is compact and has a high heat transfer rate on the tube-side. A disadvantage of the twotube pass design is that while the upper-portion of the tube has counter-current flow, the lowerportion of tube has co-current flow which limits approach (Th2 Tc2) temperatures. There also may be issues with fouling and cleaning of the tubes. The multi pass heat exchanger has higher frictional losses than multiple-tube heat exchangers.",
        "texts": [
            " In a laboratory environment, electrical heat, molten salt or sand baths may be used. In an industrial environment, high temperature steam may be used depending on process facilities. For the case of steam, cross-flow heat exchange of steam on the shell-side would be appropriate. Chapter 8 Heat Transfer and Finite-Difference Methods 563 In heat exchanger configurations, there can be more than one tube for the tube-side or a single tube can make multiple bends within itself. A singlebend in the tube gives two passes of the tube with the shell-side fluid as shown in Figure 8.4. Introduction of baffles or guides within the shell-side allows the path of the shell-side fluid flow to be lengthened. Such refinements in heat exchanger design are made to reduce the required heat exchanger area and size, and to increase heat exchanger efficiency. There are many other possible refinements in the design of heat exchangers. Two tubes that each have a bend give four passes of the tube with the shell-side fluid. Similarly, there can be more than one division in the shellside, that is, when the shell-side is divided into two-sections, it creates two shell-passes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002372_1.4029883-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002372_1.4029883-Figure3-1.png",
        "caption": "Fig. 3 Geometrical models for the hole-pattern seals. (a) SDHP seal [13]. (b) LDHP seal [15].",
        "texts": [
            " Figure 2 depicts the hole-arrangements of the SDHP and LDHP seals. For the SDHP seal, it has 2668 holes in the stator part. For the LDHP seal, it has 196 holes in the stator. The detailed geometry dimensions and operation conditions are listed in Table 1, which is in accordance with Refs. [13,15]. The commercial software ANSYS ICEMCFD 11.0 is used to generate the geometrical models and meshes for the transient CFD computations. Since the orbit of rotor is off-centered, the full circumferential (360 deg sectional) geometrical models are modeled. Figure 3 illustrates the computational models for the SDHP and LDHP seals. In the present paper, the multiblock structured grids are generated for two hole-pattern seals. The computational meshes (close-up) for the hole-pattern seals are shown in Fig. 4. Since there are round holes in the seal stator, O-type grids are generated to improve the grid quality. To simplify the modeling and mesh generation process, the feature copy, translate, rotate, mirror, and merge operations are performed. In order to resolve the flow fields in the boundary layer, meshes are refined near all the wall surfaces"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000500_s11665-012-0227-y-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000500_s11665-012-0227-y-Figure10-1.png",
        "caption": "Fig. 10 Effective stress distribution of thin-walled rectangular 3A21 tube in the whole process: (a) after bending to 90\u00b0, (b) after retracting mandrel, and (c) after springback",
        "texts": [
            " Since springback is the last step of the whole forming process, all of the errors appeared in the bending process, and the retracting mandrel process are accumulated to springback calculation stage. Therefore, the springback prediction precision depends greatly on the stress simulated in the bending process and the retracting mandrel process, which has to be studied thoroughly. The distribution and variation of the stress were obtained using the FE model established above. The distribution of the effective stress of the thin-walled rectangular 3A21 tube is shown in Fig. 10. It can be seen that retracting mandrel and springback has released a large amount of the effective stress. Some representative sections are chosen to analyze the tangential stress more clearly. Figure 11 shows the represen- tative nodes A ~ J and A\u2032 ~ J\u2032 on the edges of intrados and extrados in the bending part when the tube is bent to 90\u00b0. J-J\u2032 is the initial bending section, and A-A\u2032 is defined as a reference section which remains unchanged in the bending process. B-B\u2032 ~ I-I\u2032 are 10\u00b0~ 80\u00b0 to the reference section A-A\u2032, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000735_icra.2014.6907448-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000735_icra.2014.6907448-Figure4-1.png",
        "caption": "Fig. 4. 6-DOF CDPM example",
        "texts": [
            " For each pose of the moving platform path, the dynamic equation of a CDPM under a harmonic excitation can be written as: Mx\u0308(t)+Ke(\u03c9)x(t) = Fe j\u03c9t . (14) Assuming that the vibration response of the moving platform is x(t) = Xe j\u03c9t , solve (14): X = ( \u2212\u03c92M+Ke(\u03c9) )\u22121 F. (15) The dynamic amplification due to resonance will enable to identify the manipulator natural frequencies. In this section, a 6-DOF cable-suspended parallel manipulator driven by 6 cables is presented as an example to investigate the effects of sagging cable on the static and dynamic characteristics of CDPM. As presented in Fig. 4, there are 6 attachment points on the three vertical poles and 6 attachment points on the moving platform. These points are connected by 6 cables. This configuration is similar to the prototype presented in [1]. Table I gives the configuration parameters of the 6-DOF CDPM. The effect of sagging cable on the static behavior of CDPM is firstly considered. 1) Definition of Pose Error: For a traditional rigid-link manipulator, the platform pose error can be defined by its Cartesian stiffness matrix, assuming the compliant displacements of the platform is small [5]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure4-1.png",
        "caption": "Fig. 4. Simplified model for the deformation of the planet shaft caused by a load acting on the same shaft.",
        "texts": [
            " It is thus convenient to derive an influence coefficient from the combined deformations of the planet shaft and the carrier. The deflection of the planet shaft, which is determined according to the integration equation, \u03b4(z) = \u222b \u222b M(z) E 1 I(z) d zd z + c 1 \u00b7 z + c 2 , (23) can be solved by using the method of singularity function (see Appendix ) under different boundary conditions considering the deformation influences of the carrier. The deflection \u03b4i - i ( z ) of the i th planet shaft due to the load acting on the planet i in this case is analyzed by using the simplified model shown in Fig. 4 . Because the bending deformation of the side plates of the carrier affects the planet shaft less than the twist deformation, only the twist of the carrier is considered in the study to simplify the analysis. The planet shaft can be modeled with two elastic supports having the spring stiffness of the value k CA (near the input side) and k CB (near the output side), respectively, Fig. 4 . Either the displacement \u03b4A or \u03b4B is caused by the reaction forces F A and F B on both the supports simultaneously. More details are discussed in the next section. The bending moment M i,j along the axis, which is necessary for the deflection, is expressed with the singularity function, M i,i (z) = \u2212F A \u00b7 S(z, 0 , 1) \u2212 M A \u00b7 S(z, 0 , 1) + F i \u00b7 S(z, z BRG , 1) . (24) Either the displacement \u03b4A or \u03b4B is caused simultaneously by the force F A and F B acting on the plate A and B of the carrier, and can be expressed by using the equivalent stiffness accordingly, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001615_rpj-05-2017-0078-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001615_rpj-05-2017-0078-Figure5-1.png",
        "caption": "Figure 5 Schematic of laser scanning",
        "texts": [
            " Three additional plate-type samples with bigger areas were fabricated, and such a proportion was verified. The coating power and energy consumption held constant, which equals to 11.22, 10.5 and 9.8 per cent of total scanning energy. Therefore, it is found that varying exposure time between 60 and 80 ms does not alter the build time, but proportionally changes the scanning power, consequently, consuming different amounts of energy in SLM fabrication using AM250. Here is our explanation. As shown in Figure 5, the laser beam dwells at each point for a short while, defined as exposure time te, and the scanning power is expressed as Ps. Then, the laser deflector takes a jumping time tj to move the laser to the next point, and the jumping power is expressed as Pj. Pj is normally smaller than Ps. The time measured was actually the sum of te and tj. Though te is changed from 60 to 80 ms, the existence of tj, which is larger than te, attenuates the change of the total build time. The current sampling interval of the wattmeter is 1 s, Figure 4 Power profile during the recoating and melting Exposure time (ms) Laser scanning Coating 60 70 80 Average power (W) 1,026 1,096 1,174 380 Processing time of one layer (s) 33 33 33 10 Total energy consumption (kJ) 1,218"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003093_s1068366617030163-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003093_s1068366617030163-Figure2-1.png",
        "caption": "Fig. 2. Deviation from a cylindrical shape for the formation of several pressure zones (fluid wedges): (a) two-lobe with elliptical bore; (b) two-lobe with displaced (stepped) bore; (c) three-lobe; (d) four-lobe.",
        "texts": [
            " To decrease the level of vibrations, conditions such as multilobe bearings in which the shaft is almost hanging and so-called clearance adjustment does not occur are usually used. To form two fluid wedges, a cylindrical bearing is manufactured by the elliptical boring modeled in [1]. Before boring the working surfaces, a shim of a specific thickness is inserted between two bushes and subsequently removed during the installation of bushes. As a result, the shape of the bearing resembles the longitudinal cross section of a lemon (Fig. 2a). A two-lobe SB can be obtained according to [4] by shifting bushes in the division plane (Fig. 2b). Multilobe bearings can exist with three (Fig. 2c) and four lobes (Fig. 2d) (the techniques for manufacturing them are not described here). The multilobe design is considered in the hydrodynamic calculations by calculating and presetting the additional components of a gap. These deviations from a circular cylindrical shape refer to the designed deviations. In Fig. 3, the dimensional profiles of the additional components of gaps in two-lobe (lemon bore) and three-lobe bearings are shown. The misalignment of the shaft axis with respect to the bush axis was also modeled as a rotation of an undeformable shaft in an undeformable bush"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000735_icra.2014.6907448-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000735_icra.2014.6907448-Figure3-1.png",
        "caption": "Fig. 3. General configuration of cable-suspended parallel manipulator",
        "texts": [
            " In this case, a major control issue is to suppress the platform vibrations to ensure the positioning accuracy by decreasing the stabilization time, such as pick-and-place applications. So inherent dynamic coupling analysis between manipulator motions and dynamic cable stiffness is strongly needed for the two kinds of CDPMs. This section focuses on cable-suspended parallel manipulator but the method can be extended to non-suspended cable parallel manipulators. In this section, the dynamic stiffness matrix of cable-suspended parallel manipulator is established. The dynamic response functions and the natural frequencies are used to investigate its properties. Fig. 3 presents a general configuration of cable-suspended parallel manipulator. Ai is the attachment point in the moving platform. Bi is the attachment point in the fixed base. \u211cG (OG,x,y,z) is the fixed global frame. \u211ce (oe,xe,ye,ze) is the local frame located on the moving platform. Li is the length of the i th cable. Firstly, the dynamic stiffness matrix of each cable Ki(\u03c9) (2) should be expressed in the global frame (Fig. 3) as: KGi(\u03c9) = Tr\u22121 i Ki(\u03c9)Tri, (10) where Tri is the rotation matrix from the i th local cable frame to the global frame. Then the stiffness matrix of the manipulator KM(\u03c9) can be assembled by considering all driven cables: KM(\u03c9) = m \u2211 i=1 AT i KGi(\u03c9)Ai, (11) where: Ai =  1 0 0 0 \u2212z\u2212\u2212\u2192oeAi y\u2212\u2212\u2192oeAi 0 1 0 z\u2212\u2212\u2192oeAi 0 \u2212x\u2212\u2212\u2192oeAi 0 0 1 \u2212y\u2212\u2212\u2192oeAi x\u2212\u2212\u2192oeAi 0  . (12) Furthermore, the manipulator stiffness matrix can also be expressed as Ke(\u03c9) in the platform frame \u211ce (oe,xe,ye,ze) using the rotation matrix Te"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000637_1350650114538779-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000637_1350650114538779-Figure2-1.png",
        "caption": "Figure 2. (a) Top view of the surface consisting spherical micro-dimples; (b) geometry of a small imaginary square cell consisting of a small micro-dimple.",
        "texts": [
            " In the present study, an attempt has also been made to study the effect the direction of texturing (forward and backward) on the dimple textured slider bearing in addition to effect of inclination. A theoretical model has been developed to study the slider bearing property, for a textured surface consisting of spherical micro-dimples. Here, the slider bearing consists of two parallel plates. The clearance between two plates is h0. The upper plate is moving with a velocity U, when the lower plate is fixed as shown in Figure 1. The lower plate consists of spherical micro-dimples as shown in Figure 2(a). The dimple base radius is rp and the maximum dimple depth is hp. At any distance the height of the fluid film (local film thickness) is h (x,z). Dimensionless clearance, , dimensionless dimple depth, hp, dimensionless length x, dimensionless local length, x , dimensionless local width, z , dimensionless bearing length, L, dimensionless bearing width, B, dimensionless local height, h, and textured portion of slider width, , are given as \u00bc h0 2rp , hp \u00bc hp h0 , x \u00bc x rp , x \u00bc x rp , z \u00bc z rp , L \u00bc L rp , B \u00bc B rp , h \u00bc h h0 , \u00bc Bp B The spherical dimple textured surface is developed in the periodic repetition of a small dimple in a small imaginary square cell of arm length of 2r1 on the textured surface. It would be convenient to analyze the surface considering a small square cell as shown in Figure 2(b). Dimple area density, Sp, dimensionless half side of imaginary square cell, r, are given in equations (1) and (2) Dimple area density \u00f0Sp\u00de \u00bc projected area of the dimple in the imaginary square cell area of the imaginary square cell Sp \u00bc r2p 2r1 2r1 \u00f01\u00de r \u00bc r1 rp \u00bc 1 2 ffiffiffiffiffi Sp r \u00f02\u00de For finding out the expression of fluid film thickness (h) at any point of the textured surface, the geometry of side view of a single spherical dimple is shown in Figure 3. From Figure 4 BD \u00bc rp OD \u00bc OA \u00bc a AB \u00bc hp O0B \u00bc h0 OO0 \u00bc OA O0B\u00fe BA\u00f0 \u00de\u00bd OO0 \u00bc a h0 \u00fe hp From OBD a hp 2 \u00fer2p \u00bc a2 \u00f03\u00de at NORTH CAROLINA STATE UNIV on March 13, 2015pij"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000212_detc2013-12809-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000212_detc2013-12809-Figure2-1.png",
        "caption": "FIGURE 2. THE ASSEMBLY OF THE DISC-PAD SYSTEM.",
        "texts": [
            " This requires significantly low computational time as compared to the Lagrangian approach. Figure 1 shows schematically both the Eulerian and Lagrangian approaches. Recently, Str\u00f6mberg [20] developed and implemented an Eulerian approach for simulating frictional heating in disc-pad systems. In the present work, a toolbox developed by Str\u00f6mberg, which is based on this approach is used to perform the frictional heat analysis of a three-dimensional finite element model of the disc-pad system shown in Fig. 2. In this Eulerian approach the contact pressure is not constant, but varies at each time step taking into account the thermomechanical deformations of the disc and the pad. This updated contact pressure information is used to compute heat generation and flow to the contacting bodies at each time step. The assembly shown in Fig. 2 is a disc brake system of a heavy Volvo truck. The disc is geometrically symmetric about a plane normal to the z-axis. It is assumed that thermomechanical loads applied to the system are symmetric so only half of this assembly is considered for the simulation. The brake pad is supported by a steel plate at the back side. Some detailed geometry of the disc and back plate has been removed to simplify the model as that is not important for this analysis. As the disc is formulated in Eulerian framework, heat is transported through the mesh by convection"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000748_tie.2013.2276025-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000748_tie.2013.2276025-Figure9-1.png",
        "caption": "Fig. 9. Comparison of measured and calculated vectors of magnetic flux density at given points in the interior of magnetic circuit (see Fig. 3).",
        "texts": [
            " That is why this value was found using the experimental calibration consisting in the requirement that the calculated and measured surface temperatures of the billet are in the closest possible relations. This way leads to the value \u03b1gen = 125 Wm\u22122K\u22121, which respects all the above items. Finally, we also verified the data provided by the manufacturer of the permanent magnets VMM10, because they are characterized by a certain variance. Important are particularly their remanences Br and directions of magnetization. Basic information is given in Fig. 9 showing the accordance between measurements realized several times at the indicated points in the area of the billet by a high-precision 2-D probe and simulation carried out on the arrangement in Fig. 3 without the presence of the billet. While the model for simulation considers perfect directions of magnetization of particular permanent magnets and the same values of their remanences, the measurements show certain differences brought about by the manner of their manufacturing. It follows from the comparison between the simulations and experiment that the real remanence corresponds to the value Br = 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001184_iet-rpg.2016.0639-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001184_iet-rpg.2016.0639-Figure1-1.png",
        "caption": "Fig. 1 Test rigs (a) Experimental test rig, (b) Gear with different faults",
        "texts": [
            " Additionally, a ratio termed as fault detectability was evaluated using these measurements to comment about the efficient measuring parameter for gear fault diagnosis under fluctuating speed conditions. Along with the fluctuating profile of speed, different loads were also applied for the assessment of various measurements. A wind turbine uses a simple motor-drive-load-type arrangement [15, 16]. In a similar way, a gear test rig which imitates wind turbine system with detachable gears was used for the measurements as shown in Fig. 1a. It consists of two support bearings mounted in casing which were deep groove ball bearings, pre-lubricated and sealed at both ends. An AC motor (2.237 kW, three phase) was used to drive the test rig using a flexible coupling. The control panel of the AC motor can provide a continuously varying speed, ranging from 0 to 3000 rpm using motor control module of NV Gate application software. Vibration isolation rubber leaves were kept beneath the motor and in between the gearbox casing and its supports to reduce the vibration transmission",
            " The transducer was mounted at the bearing case of the test gear shaft and remained outside the gearbox. The gear system used society of automotive engineers (SAE) 80W-90 to lubricate by splashing. Spur gears with 20\u00b0 pressure angle and of medium carbon steel (AISI1045) were tested with faults, namely crack tooth, chipped IET Renew. Power Gener., 2017, Vol. 11 Iss. 14, pp. 1841-1849 \u00a9 The Institution of Engineering and Technology 2017 1841 tooth and missing tooth. These faults were created in the gears by using wire electron discharge machining as shown in Fig. 1b. The specifications of gears are listed in Table 1. The test gears were mounted on the main shaft for testing. To acquire the acoustic signals under the fluctuating speed conditions, a microphone was placed at ten different positions in a radius of 1 m as shown in Fig. 2a. The pattern of wind blowing cannot be predicted, but it can be assumed that it blows in somewhat sinusoidal fashion. The speed was varied sinusoidally about 750 rpm, i.e. 12.5 Hz ranging from 600 to 900 rpm as shown in Fig. 2b"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001364_s40998-017-0044-2-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001364_s40998-017-0044-2-Figure15-1.png",
        "caption": "Fig. 15 Quadrotor configuration",
        "texts": [
            " 6 Evolution of the PID control gains u2 \u00f0KP2;KD2;KI2\u00de Tables 1 and 2 present a quantitative comparative study between our approach and the designed adaptive fuzzy control law in (Cui et al. 2017) with and without presence of disturbances, respectively. The results show the superiority of our approach (the fuzzy PID control with e-modification) against the criteria IAE (Integral Absolute Error), ISE (Integral Squared Error) and IAU (Integral Absolute Input). Example 2: Stabilization of a quadrotor The quadrotor is a small Unmanned Aerial Vehicle (UAV) with four propellers actuated by DC motors mounted on the end of two perpendicular arms. A basic diagram is shown in Fig. 15. Each rotors pair of the same arm rotates in same direction: one pair rotates clockwise, while the other rotates counter clockwise. The quadrotor moves by adjusting the angular velocity of each rotor. One considers the dynamical model of quadrotor in (Bouabdallah 2007) given by _x1 \u00bc x2 _x2 \u00bc x4x6a1 x4a2Xr \u00fe b1u1 _x3 \u00bc x4 _x4 \u00bc x2x6a3 \u00fe x2a4Xr \u00fe b2u2 _x5 \u00bc x6 _x6 \u00bc x2x4a5 \u00fe b3u3 _x7 \u00bc x8 _x8 \u00bc cos x1\u00f0 \u00desin x3\u00f0 \u00decos x5\u00f0 \u00de \u00fe sin x1\u00f0 \u00desin x5\u00f0 \u00de\u00f0 \u00deb4u4 _x9 \u00bc x10 _x10 \u00bc cos x1\u00f0 \u00desin x3\u00f0 \u00decos x5\u00f0 \u00de sin x1\u00f0 \u00decos x5\u00f0 \u00de\u00f0 \u00deb4u4 _x11 \u00bc x12 _x12 \u00bc g \u00fe cos x3\u00f0 \u00decos x1\u00f0 \u00deb4u4; where x1 \u00bc / and x2 \u00bc _/ are, respectively, the roll angle and corresponding angular velocity; x3 \u00bc h and x4 \u00bc _h are, respectively, the pitch angle and corresponding angular velocity; x5 \u00bc w and x6 \u00bc _w are, respectively, the yaw angle and corresponding angular velocity; x7 \u00bc X, x9 \u00bc Y , and x11 \u00bc Z are the Cartesian position coordinates and x8 \u00bc _X, x10 \u00bc _Y , and x12 \u00bc _Z are the corresponding velocities, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002148_j.elecom.2015.02.010-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002148_j.elecom.2015.02.010-Figure3-1.png",
        "caption": "Fig. 3. A) Schematic representation of the designed fuel cell housing used for the characterization of the glucose fuel cell using a functionalized porous silicon wafer. B) Photograph of the functionalized porous silicon (3 \u00d7 3 mm) membrane wafer.",
        "texts": [
            "5 mol L\u22121 KOH), which was maintained using a peristaltic pump. The cathodic compartment is in contact to ambient air. The continuous glucose flow at a rate of Fig. 4. A) Performance of a glucose/air fuel cell using CNTs/PcCo catalyst at the cathode and C 0.8 mol L\u22121. Glucose flow rate 5 mL min\u22121 and air breathing cathode. B) Influence of the func membrane. C) Evolution of the fuel cell performance vs the TMA+ concentration. 5 mL min\u22121 diminishes substrate depletion, product accumulation, and allows constant mass transport. Fig. 3 presents a schematic representation of the cell with a photograph of the silicon membrane wafer. The performance of the abiotic fuel cell shows a typical bell shaped power profile with an open circuit voltage of 0.7 V (Fig. 4.A). This value indicates a clearly enhanced conduction of hydroxyl ions of this newmembrane leading to a reduction of both anodic and cathodic overvoltages. Glucose oxidation ismore promoted in basic solutions using abiotic catalysts due to enhanced activity of the catalyst and the availability of gluconates (deprotonated glucose), which are easier to oxidize"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001946_j.ijhydene.2015.09.143-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001946_j.ijhydene.2015.09.143-Figure1-1.png",
        "caption": "Fig. 1 e Schematic diagram a microbial fuel cell (material electrode granular graphite).",
        "texts": [
            " Electrochemical characterization of microbial fuel cells The first set of experiments consisted of the evaluation of the effect of the type of inoculum (In-EFe(III)-S, In-EFe(III)-SR, In-EMn(IV) and In-SR) on the electrochemical characteristics of theMFC-G. Each face of theMFC-Gwas characterized by separate (I and II), in series, and in parallel electric arrangements. Parallel connection of faces increased the maximum volumetric power PV-max up to 14 521, 15 825, 16 359 and 9293 mW/m3 (with In-EFe(III)-S, In-EFe(III)-SR, In-EMn(IV) and In-SR, respectively) (Fig. 1 and Table 2), compared with series connection where PV-max of 10 377, 12 778, 10 685 and 6842 mW/m3 (inocula in the same order as above). Parallel connection significantly decreased the Rint of the cells and almost doubled volumetric power. The PV-max for the MFC-G (In-EFe(III)-S, In-EFe(III)-SR, In-EMn(IV) and In-SR) when faces were atch operation. connected in series and parallel were higher than those reported in the literature [16,34,35]. Our PV-max was superior to that reported by Ortega-Martinez et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001933_b978-0-12-814862-4.00004-1-Figure4.12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001933_b978-0-12-814862-4.00004-1-Figure4.12-1.png",
        "caption": "FIG. 4.12",
        "texts": [
            " However, for this type of early glucose biosensors, a high operation potential is required to perform the amperometric measurement of hydrogen peroxide at high selectivity. Improved methods utilize artificial mediators instead of oxygen to transfer electrons between the GOx and the electrode [8]. Reduced mediators are formed and reoxidized at the electrode, providing an electrical signal to be measured. A blood glucose test is typically performed by pricking the finger to take blood, which is then applied to a disposable \u201ctest strip.\u201d Fig. 4.12 shows a typical glucose meter and a test strip. Each strip includes layers of electrodes, spacers, and immobilized enzymes assembled in a small package [12a]. Continued research and development have worked to reduce the overall size of the sensor itself and reduce the amount of blood required for an accurate measurement ( \u03bcL). The advanced glucose electrodes do no use mediators and measures direct transfer between the enzyme and the electrode. The electrode directly transfers electrons using organic conducting materials based on charge-transfer complexes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003682_lsens.2018.2880747-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003682_lsens.2018.2880747-Figure3-1.png",
        "caption": "Fig. 3. Hatching patterns of the fin walls.",
        "texts": [
            " 1) shows the recoating direction. 2) Width: The width of fin walls varies from 0.06 mm to 0.3 mm with the step size of 0.01 mm. 3) Height: The designed height of fin walls differs from 0.6 mm to 3.0 mm with the step size of 0.1 mm. Note that the height is proportional to the width in each thin wall with an aspect ratio of 0.1 in our experiments (see Fig. 1). 4) Hatching: The designed hatching patterns of fin walls, which employ the standard EOS processing path, are significantly different as the width increases. Fig. 3 shows 4 hatching patterns that are detailed as: 1) Fin 1 and 2 have two outer layer paths (or contours), two inner rectangle paths, and rotating diagonal hatching from rectangles. 2) Starting at wall 3 there are three outer layer paths and rotating diagonal hatching inside the innermost rectangle. 3) Starting at fin 15 there is one rectangular hatching. 4) Starting at fin 19 there is one thin area path. B. Image registration and edge characterization This experiment uses post-build X-ray CT images to quantify the geometric variations of each fin",
            " (1) denotes the error term in ANOVA model. D. Predictive Modeling In addition, we develop a regression model to quantify the relationship between edge roughness (i.e., the dependent variable) and the orientation, width, height, and hatching pattern (i.e., the independent variables) of each fin wall: \u03c3 = \u03b20 + \u03b21 \u00d7O + \u03b22 \u00d7W + \u03b23 \u00d7 H + \u03b24 \u00d7 Ha + \u03b25 \u00d7O \u00d7W+ \u03b26 \u00d7O \u00d7 H + \u03b27 \u00d7O \u00d7 Ha + \u03b28 \u00d7W \u00d7 H + \u03b29 \u00d7W \u00d7 Ha+ \u03b210 \u00d7 H \u00d7 Ha + \u03b5 (2) where Ha denotes the hatching pattern and is a categorical variable with four levels (defined in Fig. 3), O stands for the orientation which is also a categorical variable with three levels (defined in Fig. 1), and H and W represents the height and width of the corresponding fin wall, respectively. The R-square (R2) is utilized to statistically measure the performance of the model, and is defined as: R2 = 1 \u2212 Sum of Squareresidual Sum of Squaretot al = 1 \u2212 \u03a3i (\u03c3i \u2212 \u03c3\u0302i ) 2 \u03a3i (\u03c3i \u2212 \u03c3)2 (3) where \u03c3i is the edge roughness, \u03c3\u0302i is the predicted value of the variance, and \u03c3 is the overall average of the data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001043_s11044-017-9572-9-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001043_s11044-017-9572-9-Figure5-1.png",
        "caption": "Fig. 5 System with multiple contacts. Contacts between the feet and the ground are considered planar while contact between the hand and the wall is considered as a point contact",
        "texts": [
            " This area is much bigger when a 4-faced friction cone is used, compared to the case when 8-faced and 12-faced friction cones are used. Also, it can be seen how the required frictional forces and vertical torque modulate the effective size and shape of the foot, by making it smaller when large frictional forces and \u201cyaw friction\u201d is required. When multiple spatial contacts between the robotic system and the environment exist, it is easy to generalize condition (15). An example of such a system is given in Fig. 5, where two planar contacts exist between the ground and the feet, as well as point contact between the robot\u2019s hand and vertical wall. Contact wrenches between the ground and the feet must satisfy condition (15), while contact between the hand and vertical wall has to satisfy (3). It is important to recall that these constraints are given in local coordinate frames of the contacts. Thus, if we express contact forces in a world coordinate frame, we must pre-multiply them with corresponding rotation matrices in order to express them in a local coordinate frame of the contact",
            " Recently developed algorithm Kintinuous [37] is able to reconstruct the 3D space around the robot with resolution below 5 mm [8] in real time, using the Microsoft Kinect commodity RGB-D sensor. Having this in mind, the position and orientation of contact surfaces can be assumed to be known. If accuracy provided by Kintinuous algorithm is not sufficient, it could be further refined by force sensing together with internal joint angle sensors and kinematic model of the robot. In complex environments, such as the one shown in Fig. 5, it is extremely important to check if the intended motion is feasible before a robot initiates its execution. For example, a robot cannot lift its hand from the surface because the robot will lose dynamical balance and ultimately fall. Also, if the surface in contact with the hand has a low friction coefficient, the robot will be unable to maintain even static posture, since the hand will slide, causing the robot to fall. The total wrench that all contact wrenches create for the CoM of the system from Fig. 5 is: [ FC MC ] = [ FL + FR + FH rCoM L \u00d7 FL + ML + rCoM R \u00d7 FR + MR + rCoM H \u00d7 FH ] = GCFext. (17) The matrix GC represents contact matrix calculated for the CoM. It relates all contact wrenches acting on the body with total wrench acting on the CoM. If a robot with mass m needs to perform a motion with a desired acceleration of CoM ades C and desired rate of change of angular momentum L\u0307des C , contact forces need to counterbalance gravitational and inertial forces: [ mades C L\u0307des C ] = [ FC MC ] + [ mg 0 ] ",
            " Although, the framework is developed as a part of previous research, a short description will be given in order to make the paper clearer and self-contained. The way of including motion constraints summarized in Table 2 into task-prioritization framework will be briefly described in this section. Contacts constrain the system, which can be written as J\u0304q\u0307 = 0, where the vectors of robot\u2019s joint coordinates, velocities, and accelerations are q, q\u0307 and q\u0308, respectively. J\u0304 represents a composite Jacobian matrix for all contacts between the robot and the environment. For the case shown in Fig. 5, it has the form J\u0304 = [JT L JT R JT H lin]T , where JL and JR are Jacobian matrices for the left and right foot, while JH lin is just a linear part of the Jacobian matrix associated with the robot\u2019s hand. Dynamics of the multi-body system with contacts is given by Hq\u0308 + h0 = \u03c4 + J\u0304T Fext, (22) where H is the symmetric positive-definite inertia matrix, h0 is the vector which includes the velocity and gravitational loads, and \u03c4 is the vector of joint torques. The constraint equation J\u0304q\u0307 = 0 can be differentiated with respect to time, yielding J\u0304q\u0308 + \u02d9\u0304Jq\u0307 = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000637_1350650114538779-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000637_1350650114538779-Figure3-1.png",
        "caption": "Figure 3. Side view of a small micro-dimple.",
        "texts": [
            " It would be convenient to analyze the surface considering a small square cell as shown in Figure 2(b). Dimple area density, Sp, dimensionless half side of imaginary square cell, r, are given in equations (1) and (2) Dimple area density \u00f0Sp\u00de \u00bc projected area of the dimple in the imaginary square cell area of the imaginary square cell Sp \u00bc r2p 2r1 2r1 \u00f01\u00de r \u00bc r1 rp \u00bc 1 2 ffiffiffiffiffi Sp r \u00f02\u00de For finding out the expression of fluid film thickness (h) at any point of the textured surface, the geometry of side view of a single spherical dimple is shown in Figure 3. From Figure 4 BD \u00bc rp OD \u00bc OA \u00bc a AB \u00bc hp O0B \u00bc h0 OO0 \u00bc OA O0B\u00fe BA\u00f0 \u00de\u00bd OO0 \u00bc a h0 \u00fe hp From OBD a hp 2 \u00fer2p \u00bc a2 \u00f03\u00de at NORTH CAROLINA STATE UNIV on March 13, 2015pij.sagepub.comDownloaded from a \u00bc r2p \u00fe h2p 2hp \u00f04\u00de OO0 \u00bc r2p \u00fe h2p 2hp h0 \u00fe hp \" # \u00f05\u00de Now the equation of sphere of radius a is x2 \u00fe y2 \u00fe z2 \u00bc a2 \u00f06\u00de By shifting the coordinate system to the edge (O0) of the imaginary square cell, the following equation has been obtained y \u00bc h0 \u00fe hp a\u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a2 x r1\u00f0 \u00de 2 z r1\u00f0 \u00de 2 q \u00f07\u00de at NORTH CAROLINA STATE UNIV on March 13, 2015pij"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001958_s12555-014-0472-y-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001958_s12555-014-0472-y-Figure1-1.png",
        "caption": "Fig. 1. Two inverted pendulums connected by a spring and a damper.",
        "texts": [
            " sj j j j j p e p e \u03c8 \u03c8 \u0398 = \u2212 \u2212 \u2212 +F F The remaining design procedures of the virtual controller, adaptive laws and control law are similar to those in the asymmetric constraint case and are omitted. 5. APPLICATION EXAMPLES 5.1. Simulation example In this section, a simulation of a nonlinear MIMO system for evaluating the performance of the prescribed error constraints against uncertainties is presented. As an application system, two inverted pendulum systems composed of a spring-and-damper connection and nonlinear friction as shown in Fig. 1, are used [17]. The pendulum angle and angular velocity are controlled by the torque inputs generated by a servomotor at each base. The two-inverted-pendulum system has the following dynamic equations: 1 2 1 1 1 1 1 1 2 2 2 2 2 2 sin 0.5 cos( ) , sin 0.5 cos( ) , f f J m gl Fl T u J m gl Fl T u \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 = \u2212 \u2212 \u2212 + = + \u2212 \u2212 + (68) where \u03b81 and \u03b82 are angular positions, 1 0.5J = kgm2 and 2 J = 0.625 kgm2 are the moments of inertia, 1 2m = kg and 2 2.5m = kg are the masses, 0.5l = m, i u are Seong-Ik Han and Jang-Myung Lee 8 control inputs, Tfi are the friction torques, ( )F k x d= \u2212 bx+ denotes the force applied by the spring and damper at the connection points, and x is the distance between the connection points A and B as follows: 2 2 1 2 2 1 (sin sin ) [1 cos( )], 2 l x d dl \u03b8 \u03b8 \u03b8 \u03b8= + \u2212 + \u2212 \u2212 where 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003966_iciset.2018.8745656-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003966_iciset.2018.8745656-Figure4-1.png",
        "caption": "Fig. 4. Circuit diagram of the proposed system.",
        "texts": [],
        "surrounding_texts": [
            "This process is handled by the Google speech recognition engine. But at first, need to install the python. After that may software module will be needed. So, that using a package manager is a right decision. We are using pip as package manager of python. For manipulating voice we need something. PyAduio helps us in this part. Receiving voice command with a microphone and manipulating the voice is handled by the PyAudio. When we receive raw voice there can be any noise or unwanted sound element. That\u2019s why we are using a module name noise ambient, which will remove the unwanted noise in our voice command. This module helps to increase the voice command accuracy rate. We are working on a Jessie operating system. In this type of operating system when any voice command received and send to somewhere, the main voice command can be lost or corrupted. For that, we need to use the FLAC encoder. This FLAC encoder will encode our voice command and make it stable. This will not be needed if anyone working in windows, Linux or Mac operating system. At first, the voice is captured by a microphone then it will send the voice is received by the raspberry pi. The raspberry pi sends the voice command to the google speech recognition engine. The STT (Speech to text) engine change the voice command into the text message. We manipulating this text in our system. This system can be configurable with 119 different languages. In our case, we were implementing two languages. For configuring the language we need to change the parameter of the language. There is a fixed data sheet for that. But sometimes it causes an encoding problem. We solved our encoding problem for Bangla speech recognition using perfect encoding for the raspberry pi."
        ]
    },
    {
        "image_filename": "designv11_13_0002740_b17506-6-Figure5.62-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002740_b17506-6-Figure5.62-1.png",
        "caption": "FIGURE 5.62 (a) Voltage-limiting ellipses and current-limiting circles for IPM machines and (b) overlap area between voltage limiting ellipses and current-limiting circles.",
        "texts": [
            " The center of the ellipse is located at (\u2212(\u03c8 m/Ld), 0), which is the characteristic current of the machine. The eccentricity of the ellipse can be represented as e b a b e V L V L V L L L s d s q s d d q = \u2212 = \u2212 = \u2212 \u239b \u239d\u239c \u239e \u23a0\u239f 2 2 2 2 2 1 ( ) ( )max max max \u03c9 \u03c9 \u03c9 (5.60) As shown in Figure 5.61b, the ellipse shrinks inversely with the rotor speed. The shape of the ellipse depends on the saliency ratio. In addition, the d-and q-axis currents must satisfy: i i Isd sq s 2 2 2+ \u2264 max (5.61) Equation 5.61 represents a current circle centered at the origin with a radius of Is max as shown in Figure 5.62a. Unlike the voltage-limiting ellipses, current-limiting circles remain constant for any speed. Since both Equations 5.57 and 5.61 should be satisfied during the operation, for a given rotor D ow nl oa de d by [ C or ne ll U ni ve rs ity ] at 1 0: 56 0 3 A ug us t 2 01 6 speed, the current vector can be located anywhere inside or on the boundary of the overlap area between the voltage-limiting ellipse and current-limiting circle as shown in Figure 5.62b. The overlap area becomes smaller when the rotor speed keeps increasing indicating progressively smaller changes for current vector in the flux-weakening region. Below base speed, where the phase voltage is less than Vs max for the rated current, the operation of the IPM machine is based on the control of q-axis current as shown in Figure 5.63a. In the fluxweakening region, negative d-axis current is applied as shown in Figure 5.63b. This creates a negative voltage vector on the d-axis, which opposes the one induced by the PM flux linkage, so that the motor can speed up"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003416_15325008.2018.1444689-Figure17-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003416_15325008.2018.1444689-Figure17-1.png",
        "caption": "FIGURE 17. Fault current with two phases open (Vc &Vd).",
        "texts": [],
        "surrounding_texts": [
            "In this paper, a dynamic model of the wind-driven FPIG in self-excited mode is developed using the two-axis d-q equivalent circuit. The simulation results for five-phase SEIG under varying wind velocities and different loading conditions are presented. The performance of five-phase wind electric generator is compared with three-phase generator in terms of voltage, current, and output power. It is found that the generator output is increased considerably for the same wind velocity with reduced per phase current. The effect of modulation indices is analyzed and it is noted that optimum modulation index varies for different wind speeds. Further the reliability of the WECS is increased enabling the generator to operate with one or two phases open. The above mentioned advantages can compensate for the additional installation costs in standalone applications where reliability is a main concern. ORCID Sanjeevikumar Padmanaban http://orcid.org/0000-00033212-2750"
        ]
    },
    {
        "image_filename": "designv11_13_0000782_j.ifacol.2015.06.433-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000782_j.ifacol.2015.06.433-Figure4-1.png",
        "caption": "Fig. 4. Torch holder",
        "texts": [
            " Ottawa, Canada shielding gas or flow, proportionality in the mixture of shielding gases and following the deposition of layers. 3. Relation study between arc welding parameters and weld bead geometry in individual and multi-layered deposition 4. Development of process control algorithms based on this study. 5. Evaluation of the behavior and stability of the process; 6. To investigate the suitability of the welding process used for depositing other alloys. The controlled mechanical device named torch holder (Figure 4) proposed herein has been developed in order to control the speed and position on translation of XY axis displacement for deposition of welding layers. Some features of the proposed system are presented as high control of the basic parameters for process\u2019 stability when it\u2019s compared to existing technologies for welding alloys. The Figure 4 illustrates the torch holder proposed: The construction of a Manufacturing Additive component has several issues to determine the success of the procedure [14]. Deposition of the first string is among the factors that require careful consideration. For this first string, it\u2019s looked for a good metallurgical bond with the substrate and a suitable thickness and morphology. On tests for a \"wall\" formation by deposition of multiple layers, the morphology and microstructure of deposition are analysed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001564_s00170-018-2582-9-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001564_s00170-018-2582-9-Figure2-1.png",
        "caption": "Fig. 2 Distribution of milling forces",
        "texts": [
            " x(t) and y(t) are the current amplitude of the vibration at the tool tip, and x(t-T) and y(t-T) are the previous tooth period amplitude of the vibration at the tool tip in the X and Y directions, respectively. The dynamic chip thickness is represented by h, as it is depicted in Fig. 1. It is a function of the immersion angle \u03c6 and it can be formulated as (x(t) \u2212 x(tT))sin(\u03c6j) + (y(t) \u2212 y(t-T))cos(\u03c6j). The machining force consists of two components: the cutting force (Fc) and the process damping (Fpd) force [8]. It is expressed as Eq. 3. The forces at the tool tip are shown in Fig. 2. F \u00bc Fc \u00fe Fpd \u00f03\u00de Fct and Fcr are the cutting forces in the tangential and radial directions, respectively. Fpdr and Fpdt are the process damping forces in the tangential and radial directions, respectively. \u03a6 is the immersion angle (rad), \u03b2f is the angle between the resultant process damping force and the tangential process damping force (rad), and \u03b8 is the angle between the resultant cutting force and the tangential cutting force (rad). The cutting forces have been projected in the x and y directions without consideration of the effect of the helix angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002720_978-3-319-24055-8-Figure5.2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002720_978-3-319-24055-8-Figure5.2-1.png",
        "caption": "Fig. 5.2 Studied gears for STE computation and optimization",
        "texts": [
            " the amount of material remove on the teeth tip, \u2022 the starting tip relief diameters / of pinion and driven gear, \u2022 the added up crowning centered on the active tooth width Cb;i=j. The optimization of tooth modifications in simple mesh gear system for a given torque has been studied by many authors [4\u20136] but the approach for multi-mesh gear systems optimization is still unusual [7]. The first part of this paper presents a complete optimization process for a truck timing cascade of gears displayed in Fig. 5.2. In this study, the first cascade is designed with three helical gears and has 8 optimization parameters (2 by gear, and 1 by mesh). The second cascade is designed with two gears and has therefore 5 optimization parameters. Moreover the modifications made on teeth profile have to be satisfying for a wide torque range. That requires an efficient method as the number of possible solutions is extremely large, due to the combinatory explosion phenomenon. The Particle Swarm Optimization (PSO) [8] has been chosen because it is particularly efficient as it is an order 0 meta-heuristic, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003599_j.procir.2018.04.033-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003599_j.procir.2018.04.033-Figure6-1.png",
        "caption": "Fig. 6 Experimentation specimens; a) massive and b) lattice structure",
        "texts": [
            " In conclusion, on these defect types based on the bibliography sources, geometry and dimensional variations correspond to the measure of tolerances of the simple feature manufactured by additive manufacturing. The tolerance of part depends on effect of staircase and effect of the machine positional error (source energy and direction of the platform motion). From the authors\u2019 experiences, Piaget et al. [24] have identified a correlation between the geometry defect of specimens and their position in the manufacturing space. Therefore, 25 locations of the manufacturing space were tested in the experimentations. Two kinds of test specimens have been chosen for the test, see Fig. 6. The massive specimen is used to measure the height variation. The lattice structure is used to observe the variation of the geometry of the cell. The authors have obtained different kinds of defects and they concentrate their study on the distortion of the first layers, see Fig. 7. The distortions of the first layers of 25 massive specimens were measured by a three-dimensional optical control machine (Vertex from Micro-vu). This equipment is able to measure the height of each point in the first layer, see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000510_s1000-9361(11)60417-2-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000510_s1000-9361(11)60417-2-Figure5-1.png",
        "caption": "Fig. 5 Structure of turbo-shaft engine.",
        "texts": [
            " Through a series of equilibrium computations and calculations of dynamic equations, such as balancing rotor model and fuselage model, the required power of helicopter can be worked out and sent to the turbo-shaft engine together with some flight parameters. In Ref. [21], a great number of tests were performed to check the accuracy of this helicopter model. The results prove that the helicopter model has a good reliability and is able to carry out digital flight simulations of routine missions. 2) Turbo-shaft engine model Figure 5 depicts a structure of the turbo-shaft engine component-level model, which is established based on T700 engine data set. In Fig. 5, the engine station numbers represent the inlet or outlet of engine components. During engine modeling procedure, every component model of the engine is created using engine thermodynamic characteristic and typical experimental data at first. And then, with the benefit of power balance, flow equilibrium, pressure equilibrium and rotor dynamics equations, the balance equations among engine components can be constructed and calculated. Finally, Newton-Raphson method and one-passthrough algorithm are adopted in order to solve steady-state engine model and dynamic engine model respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001146_0954406217722380-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001146_0954406217722380-Figure5-1.png",
        "caption": "Figure 5. Applied forces on the cage and its degrees of freedom.",
        "texts": [
            " So the time-varying reaction forces of support B1 (Fy1, Fz1) and B2 (Fy2, Fz2) due to unbalance are obtained as follows Fy1 \u00bc k1y\u00f0 yr a rz\u00de Fz1 \u00bc k1z\u00f0zr \u00fe a ry\u00de Fy2 \u00bc k2y\u00f0 yr \u00fe b rz\u00de Fz2 \u00bc k2z\u00f0zr b ry\u00de \u00f02\u00de The ball bearing and support B1 are assumed to be in static equilibrium illustrated in Figure 4(b), so the forces of bearing in the support B1 are calculated as follows Fy \u00bc Fy1 Fz \u00bc Fz1 \u00f03\u00de Dynamic model of bearing cage A dynamic model of the bearing cage with threeDOFs is developed, in which the interactions among the balls, inner rings and cage are taken into account, and the reaction forces due to rotor unbalance are also introduced as boundary conditions of bearing. 1. Dynamic equations of cage The cage motions are represented by its mass center oc in the inertial frame oyz described by the vector (yc, zc) and its rotational angle c. The forces applied on the cage and its degrees of freedom are illustrated in Figure 5. mj is the rotational angle of the jth ball. The normal contact force and friction force between cage and ball are denoted by Fcbj and fcbj, which can be decomposed to Fcbjy, Fcbjz, fcbjy, fcbjz in y and z directions respectively. Fci and Mc are the interaction force and moment between cage and inner ring. The interaction forces are denoted as F 0ciy and F 0ciz, which are transformed to the components in the inertial frame to form Fciy and Fciz in y and z directions, respectively. The force of cage unbalance is denoted by Fmc, which can be decomposed to Fmcy, Fmcz in y and z directions. The cage motion equations can be built according to the force components in the inertial frame oyz illustrated in Figure 5. Fciy \u00fe XN j\u00bc1 \u00f0Fcbjy \u00fe fcbjy\u00de \u00fe Fmcy \u00bc mc \u20acyc Fciz \u00fe XN j\u00bc1 \u00f0Fcbjz \u00fe fcbjz\u00de \u00fe Fmcz \u00bc mc \u20aczc Mc \u00fe Xj\u00bcN j\u00bc1 Fcbjy dm 2 \u00bc Ic \u20ac c \u00f04\u00de where mc and Ic are the mass and rotational inertia of cage, respectively. Forces of inner ring on the cage. The interaction force Fci and moment Mc in equation (4) can be solved by using the hydrodynamic solution in Ye and Wang10 and Yan et al.34 Force of cage unbalance. The cage unbalance force in equation (4) can be expressed as follows Fmc \u00bc 0:5mecdm! 2 c \u00f05\u00de where mec is the cage unbalance mass"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002323_1350650114545140-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002323_1350650114545140-Figure1-1.png",
        "caption": "Figure 1. Schematics of piston\u2013cylinder system and its free body: (a) piston\u2013liner; (b) schematic of film bearing; (c) forces acting on the piston; (d) vibration model of liner.",
        "texts": [
            " The squeeze item in equation (1) can be further expanded to obtain the transverse velocity of the piston, that is @h @t \u00bc _et t\u00f0 \u00de\u00fe _eb t\u00f0 \u00de _et t\u00f0 \u00de\u00bd y L \u00fe _st t\u00f0 \u00de n o cos \u00fe @ d ,x, t\u00f0 \u00de\u00bd @t \u00f03\u00de To solve equation (1) for the film pressure, the following pressure boundary conditions are employed ph , 0\u00f0 \u00de \u00bc ph ,L\u00f0 \u00de \u00bc 0 \u00f04a\u00de at Purdue University on September 1, 2014pij.sagepub.comDownloaded from ph , y\u00f0 \u00de \u00bc 0, 4 4 or \u00fe 4 42 \u00f04b\u00de @ph @ \u00bc0j \u00bc @ph @ \u00bc j \u00bc 0 \u00f04c\u00de Here, is the bearing angle of thrust side (TS) and anti-thrust side (ATS) as shown in Figure 1(b). By the use of ph obtained, the total normal force and its moment due to the hydrodynamic action of the lubricant can be separately computed by the following expressions Fh \u00bc R Z Z A ph\u00f0 , y\u00de cos d dy \u00f05\u00de Mh \u00bc R Z Z A ph , y\u00f0 \u00de a y\u00f0 \u00de cos d dy \u00f06\u00de where A represents the total bearing area of the piston skirt with radius R. The shear stress due to the hydrodynamic effect can be written as h , y\u00f0 \u00de \u00bc U h f \u00fe fs \u00fe fp h 2 @ph @y \u00f07\u00de where f, fs, and fp are the shear stress factors, whose expressions can be found in the study of Patir and Cheng"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002067_s11249-013-0291-y-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002067_s11249-013-0291-y-Figure5-1.png",
        "caption": "Fig. 5 Structure of the calibration testing bench on the electromagnetic loading and friction torque sensors",
        "texts": [
            " (3) result in Fe \u00bc 15:83 106 I2 c d2 c \u00f04\u00de As can be seen from Eq. (4), the EMLD loading depends on the working gap between the EMLD and electromagnetic cover. Specifically, the electromagnetic force increases as the gap decreases. In this study, the design working gap dc is 0.35 mm. Prior to the experiment, the exact relationship curve between the current and electromagnetic force must be obtained through calibrating the EMLD. A calibration testing bench has also been designed based on the measurement, and its structure is shown in Fig. 5. As shown in Fig. 5, the gap between the electromagnetic cover and flange can be adjusted by a gap adjustment screw. The electromagnetic force can be measured by the force sensor when the current is modulated. Figure 6 presents a comparison between the experimental calibration and theoretical data when the working gap dc is 0.35 mm. The testing device of the friction torque at seal face is shown in Fig. 7 together with the EMLD. As shown in Fig. 7, the friction force generated by the mating faces can be measured by the friction torque sensor unit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003126_b978-0-12-803581-8.10316-9-Figure21-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003126_b978-0-12-803581-8.10316-9-Figure21-1.png",
        "caption": "Fig. 21 Schematic of surface plasmon resonance (SPR) spectroscopy. Pic: elte.prompt.hu.",
        "texts": [
            " Response in resonance units (RUs) versus time gives sensorgram. A background response will also be generated due to the difference in the refractive indices of the running and sample buffers. This background response is subtracted from the sensorgram to obtain the actual binding response. The background response is obtained by injecting the analyte through a reference flow cell which has no ligand or has an irrelevant ligand immobilized onto the sensor surface. The schematic approach for SPR spectroscopy is shown in Fig. 21. Surface binding interactions like small molecule adsorption [102], protein adsorption on self-assembled monolayers, antibody\u2013antigen binding, DNA and RNA hybridization [129,130], protein\u2013DNA interactions [125], binding kinetics, affinity constants, equilibrium constants, as well as receptor\u2013ligand interactions in immunosensing [124,131] can be studied using SPR biosensors. In EC-SPR there is a combination of electrochemistry and SPR. The thin metal film on the substrate serves not only to excite surface plasmons, but also acts as a working electrode for electrochemical detection or control [132]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002746_978-3-319-44156-6_8-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002746_978-3-319-44156-6_8-Figure1-1.png",
        "caption": "Fig. 1 Construction of a variable-DOF 7R mechanism using two Bennett mechanisms. a 2-DOF double-Bennett mode. b 1-DOF 7R mode",
        "texts": [
            "1007/978-3-319-44156-6_8 71 seven dimensional kinematic image space, where the coordinates are the so called Study parameters closely related to dual quaternions. In this paper, we first present the construction of a new variable-DOF 7R mecha- nism and then discuss its algebraic analysis in detail to identify all the motion modes and transition configuration of the 7R mechanism. In [5], 1-DOF multi-mode 7R mechanisms were constructed by combining two over- constrained 4R, 5R and 6R mechanisms. In this paper, a new variable-DOF 7R mech- anism is constructed by combining two Bennett mechanisms in the following way (see Fig. 1a). First, arrange two Bennett mechanisms [1] with revolute (R) joint axes b1, b2, b3, b4 and b\u20321, b \u2032 2, b \u2032 3, b \u2032 4 in such a way that b1 and b\u20324 coincide and the common normal of b4 and b1 as well as the common normal of b\u20324 and b\u20321 are aligned with the X-axis of the base frame. Then remove joint b1 of Bennett mechanism b1b2b3b4 and joint b\u20324 of b\u20321b \u2032 2b \u2032 3b \u2032 4. Finally, connect link b1b2 to b\u20323b \u2032 4 using an R joint with its axis coinciding with the axes of the original joints b\u20324 and b1 to form a 7R closedloop mechanism",
            " Therefore we have to take a closer look at the solutions of equations E1, E2 and P1 or P2, respectively when discussing the motion modes associated with these two polynomial equations (see Sects. 3.1 and 3.2 for details). Although one could use a re-parameterization that maps the parameter value infinity to a finite one and compute everything again, the process would be more tedious. As one can see from Fig. 2b, the red curve given by P1(v1, v2) = 0 in Eq. (12) corresponds to two 1-DOF 7R modes. In a 1-DOF 7R motion mode (Fig. 1b), the remaining joint variables can be solved by first computing a Groebner basis of the polynomials E1, E2 and P1 that yields a linear polynomial in v5. Back substitution into \ud835\udc29 (the solution of the linear system) yields the pose of the coordinate frame attached to the fourth link. Together with P1 = 0 it is possible to compute this pose \ud835\udc29a(v1) depending on the input angle v1 of the 7R-chain, which is too long to be displayed here. Substitution of this pose into the remaining hyperplane equations Hi(v1, v3), Hi(v1, v4), Hi(v6) and Hi(v7) yields linear equations in the remaining unknown tangent half of the joint angles",
            " This yields the following four special configurations v1 = \u221e, v2 = 0, v3 = \u221e, v4 = \u221e, v5 = 0, v6 = \u221e, v7 = 0 (Fig. 3a), v1 = \u221e, v2 = 0, v3 = \u221e, v4 = 0, v5 = \u221e, v6 = 0, v7 = \u221e (Fig. 3b), v1 = 0, v2 = \u221e, v3 = 0, v4 = 0, v5 = 0, v6 = \u221e, v7 = 0 (Fig. 3c), v1 = 0, v2 = \u221e, v3 = 0, v4 = \u221e, v5 = \u221e, v6 = 0, v7 = \u221e (Fig. 3d). (13) As it will be shown later, through these configurations, the 7R mechanism can switch from the 1-DOF 7R modes to the 2-DOF double-Bennett mode. In this section, we will reveal that P2(v1, v2) = 0 in Eq. (12) corresponds to a 2-DOF double-Bennett mode (Fig. 1a). Solving the linear equation P2(v1, v2) = 0 with respect to v2, and substitution into \ud835\udc29 we obtain \ud835\udc29b = ( \u2212( \u221a 2 \u2212 1)v5, \u2212( \u221a 2 \u2212 1)v5, 1, \u22121, \u2212( \u221a 2 \u2212 2)v5, ( \u221a 2 \u2212 2)v5, \u221a 2, \u221a 2 )T . (14) It is noted that this point is independent of the input joint angle v1. Because this coordinate system is attached to the fourth link, which is a part of the second Bennett mechanism by construction, this link can only be driven by an additional input attached to one of the joints 5, 6 or 7. Therefore, P2(v1, v2) = 0 in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002067_s11249-013-0291-y-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002067_s11249-013-0291-y-Figure12-1.png",
        "caption": "Fig. 12 Schematics of the grooved seal plate used in the experiment",
        "texts": [
            " It is expected that the optimal sensitivity coefficient and the weight coefficient of the controllable seal will be obtained through such series of experiments. Table 1 Calibration data between the strain signal and force Static balance position controlled parameter Electromagnetic loading device\u03a3 Control unit Testing seal Temperature signal (Ti) and friction torque sensor signal (Nfi) Signal Process SystKi (Nfi , Ti) K* (Nf, T) em Adjust current + - The structure of the rotor seal used in the experiment is shown in Fig. 12. The groove pattern of the seal face is designed as consisting of a seal dam, external spiral grooves, and internal spiral grooves whose directions are consistent with the direction of rotor rotation. The seal dam separates the external spiral grooves\u2019 structure and the internal spiral grooves\u2019 structure. The external spiral grooves are designed to generate a pump-priming effect to enhance the load capacity and stiffness of the film, and easily fall into the noncontacting status under low running speed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000969_j.jsv.2017.01.018-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000969_j.jsv.2017.01.018-Figure2-1.png",
        "caption": "Fig. 2. 2DOF link model of wiper blade. (a) Upright state. (b) Buckled state.",
        "texts": [
            " Furthermore, we conducted experiments and confirmed that the result is in good agreement with an experimental observation using an actual wiper blade. Please cite this article as: M. Unno, et al., Analysis of the behavior of a wiper blade around the reversal in consideration of dynamic and static friction, Journal of Sound and Vibration (2017), http://dx.doi.org/10.1016/j.jsv.2017.01.018i A wiper blade (Fig. 1) consists of a head, a narrow neck, and a lip located at the lower part of the blade. We introduce a 2DOF analytical link model (Fig. 2(a)), which is widely used for analyzing the dynamics of a wiper blade [8,10,11]. The upper and lower links correspond to the neck and lip, respectively. The head, located above the upper link reciprocates against a swept surface fixed horizontally. Whereas the horizontal displacement of the head is predetermined, it is free moving in the vertical direction. We denote the mass of the head by m. For the neck and the lip, we denote the total masses of each by m1 and m2, their lengths by l1 and l2, and their moments of inertia at their centers of gravity by I1 and I2, respectively",
            " k0 is the rigidity of the support side. ls is the equivalent natural length of the spring k0. 2.2. Equations of motion of 2DOF link model The origin of the static \u2212x y coordinate system is established as the position of the head when the blade is standing Please cite this article as: M. Unno, et al., Analysis of the behavior of a wiper blade around the reversal in consideration of dynamic and static friction, Journal of Sound and Vibration (2017), http://dx.doi.org/10.1016/j.jsv.2017.01.018i upright (Fig. 2(a)). The various displacements of parts of the system are depicted in Fig. 2(b); specifically, ( )u v, , ( )u v,1 1 , ( )u v,2 2 and ( )u v,3 3 indicate the displacements of the head, of the center of gravity of the neck, of the center of gravity of the lip and of the lip tip, respectively. The rotational angles of the neck and the lip are denoted by \u03b8 and \u03c6, respectively. The initial compressive force applied to the wiper blade is represented by pressing the support side down, where the displacement of the support side is hd. We consider the case when the lip tip is in contact with the swept surface at all times"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001188_lra.2017.2749280-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001188_lra.2017.2749280-Figure2-1.png",
        "caption": "Fig. 2. The geometry of an example compliant part with a 1D parameterized geometry. The red line shows the geometry of the underlying finite-element model. The part thickness is halved in the areas highlighted with the blue circles, from 2 mm to 1 mm.",
        "texts": [
            " One of the main considerations, therefore, is for the system to generate plans for the system that will minimize the number of gripping relocations. The overall plan generated for the robot system depends on the predictions from the part deformation model. The same deformation model as in [1] is used in this letter. A full description of the model structure is given in [1] and only a brief overview is provided here. For each part, an overall one-dimensional geometry is defined, as exemplified by the part shown in Fig. 2. This geometry is approximated by a mesh of Ne one-dimensional elements. A single element ej is connected to two nodes, which each have two degrees-of-freedom (DOF): vertical displacement and rotation. Because only Cartesian forces are considered to be applied by the robot in our application, a torsion component is not included with the corresponding additional degrees-offreedom in the model. The values of the 2(Ne + 1) DOFs for the Ne + 1 nodes are collected in a global displacement vector V \u2208 R2(Ne +1) , which fully defines the state of the mesh",
            " For example, in the example comparison, only three samples include the very first element, and all three of those involve 12 or more elements in the effective stiffness measurement. For the incremental method, this leads to almost no change in the first element estimate. However, note that using the same data, the batch estimation algorithm obtained a much more accurate estimate of this element. For elements in the center of the mesh, where many more samples are relevant, the accuracy is quite good, even with relatively few samples obtained. Physical validation of the model parameter estimation was performed on an example thin plastic part, with the geometry shown in Fig. 2. The part was uniformly thick in all places except for two short segments on the part arms where the part thickness was halved. The overall effect was to greatly reduce the effective stiffness of the part when the thin region was between the grasping and force application points, effectively disallowing the ability to clean stains in such a configuration. The stains were manually created with a highly visible color marker to aid perception and required a nontrivial amount of normal force (1.0\u20132"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000633_icra.2014.6906622-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000633_icra.2014.6906622-Figure5-1.png",
        "caption": "Fig. 5. Geometric parameters representing tool configuration and boundary for the PQ computation proposed in [3].",
        "texts": [
            " The algorithm calculates the shortest distance between constraint points and a generalized cylinder which represents the tool shaft. For each point in a cloud (pci ), representing the constraint surface the algorithm determines the closest point on the tool shaft (pti). The tool shaft is represented by multiple generalized cylindrical segments. Each segment is enclosed by two adjacent contours. The contour center is located at pi with the normal direction ni. Scalar ri is the contour radius, as shown in Fig. 5. The positions of pc and pt are sufficient to enforce the constraint for the entire tool shaft. In the existing literature, point clouds become a common representation of object boundaries, to constraint the motion of the tool tip, due to their geometric simplicity, [17], [18]. Using the in Sec. III-B described constraint evaluation method we are able to adapt in situ to flexible constraints. Furthermore, we address an issue regarding discrete constraints which becomes important when constraining the tool shaft"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001725_iros.2018.8593566-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001725_iros.2018.8593566-Figure1-1.png",
        "caption": "Fig. 1: IRB2400 and YuMi used for the skill transfer.",
        "texts": [
            " We propose now an extended ontology for skill representation together with an implementation extension of the above mentioned interface for exporting and importing programmed skills to and from our robot-agnostic knowledge base. This implementation provides also options for the specification of synchronization points, synchronized motions and master-slave relations during program parts, which can then, based on our parameterizable skill representation, be re\u2013used and adapted into high\u2013level, robot\u2013agnostic instructions, supporting also the transfer of skills between different robots (e.g., those illustrated in Fig. 1). We illustrate the benefits of this extension and resulting capabilities in several experiments and proof-of-concept evaluations and discuss, how our approach and interface can be adapted to support the coordination between user and robot, in addition to the coordination of two robot arms, by using the same synchronization primitives. The remainder of the paper is organized as follows. We will refer to related work in the area of knowledge- or ontology-based (dual-arm) robot programming in section II",
            " In the following we describe a previously not reported proof-ofconcept experiment in which we can show the possibility of transferring skills not only between two robot arms of the same type but between entirely different robots. We performed a proof-of-concept experiment for the transfer of skills from one robot to another, in which we connected our interface to two different robots, one YuMi as previously used, and one ABB IRB2400 6DOF manipulator (due to hardware problems we used a simulation for both robots, see Fig. 1, but we could verify the general compatibility of the interface also with the physical IRB2400 in other tests). As the latter robot is not equipped with the same type of gripper as YuMi, we tested only the transfer of a simple sequence of moves in relation to two objects. We set up the test on YuMi, by defining two reference frames Pos1 and Pos2, as shown in Fig. 7, by having the right arm point to two randomly chosen positions in the robot\u2019s workspace. Fig. 8 illustrates how the interface only shows options for primitive actions available on the IRB2400 it is connected to, and gives a warning regarding the imported skill MovesWObjs needing reference objects to be available"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000796_s1068371215070111-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000796_s1068371215070111-Figure3-1.png",
        "caption": "Fig. 3. LACH of the investigated EMS with EC.",
        "texts": [
            " 7 2015 PYATIBRATOV Research on the efficiency of the electric drive damping properties shall be carried out in the terms of the EMS with EC having the following parameters of the invariable part of the system: Tconverter = 0.005, \u03c4 = 0.003 s, kelastic = 8.2, Telastic = 0.132 s, Ts = 0.0134 s, Td = 0.005 s, Tm = 0.38 s, kconverter = kengine = km = kfc = kfv = 1, and Tfc = Tfv = 0. To determine the conditions providing the best damping of torque oscillations in the elastic element Melastic.ext by the electric drive, let us consider frequency characteristics of the transmission of branches Felast.m and FED of the graph given in Fig. 2b. In Fig. 3, LACH of branch Felast.m taking into account the SMP properties, LACH LED of the back branch taking into account the ED, and LACH Lr showing change of torque in the elastic element Melastic.ext are given. Index numbers i = 1\u20134 of characteristics LED and Lr correspond to the variants of implementa tion of the SRS represented in Table 2. LACH Lr.t. shows the change of the torque of the elastic element Melastic.ext in the SMP. The research showed that, in this case, the torque amplitude in the elastic element at the resonance frequency Melastic.ext.m (\u03c9ext.m) = Melastic(\u03c9r.m)/Text(\u03c9r.m) = 9.2. The efficiency of damping of the oscillations in the mechanisms\u2019 elastic elements by the electric drive depends on certain characteristic points A, B, and C in Fig. 3. LACH LED and Lelastic.m intersect. To compare the efficiency of damping of the oscillations in the mechanisms\u2019 elastic elements by the electric drive during application of VR P and VC PI, their parame ters are selected so that, given equal values of amplifi cation coefficients VR kVR1 = kVR2 = 33.3, LACH and intersect on the left branch of LACH Lelas.m at point B. The results of modeling of the studied EMS with EC on a personal computer showed that, in the consid ered case, while using the proportional VR, the torque in the elastic element will be equal to Melastic"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002722_012101-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002722_012101-Figure4-1.png",
        "caption": "Figure 4. Joints needed for obtaining the proper operation of the model.",
        "texts": [
            " The prepared model of the whole system of the platform consists of seven components (links):  Base - which consists of a foundation and attached to it bases with eyes to which the platform is attached;  Cross - the lower joint component, which allows for the proper connection of the parallel robot leg with the base;  Engine - which contains casing of an engine with all elements of a frame that allow for correct fitting of the engine;  Screw - which consists of a screw with a frame and an end of a ball-and-socket joint;  Frame - which consists of a frame, to which is fixed a car body and 6 slots of ball-and-socket joints;  Dr_sht - a drive shaft, which ensures the transfer of power, from the engine to the nut of the screw, by means of a transmissions with a toothed belt;  Nut - nut that allows transferring a torque on the screw of a screw transmission. The next step of the model developing was determination of joints (figure 4) between the particular components of the model to reconstruct the operation of the real object with the virtual model. For the implementation of the motion simulations the following types of joints were created [11,12]:  Base_fix - joint of the fixed type that removes six degrees of freedom of the Base component,  Ki_Base - joint of the revolute joint type allowing only for rotation of the Cross component around the chosen axis,  Ki_Mi - joint of the revolute joint type allowing only for rotation of the Engine component around the chosen axis,  Mi_Wni - joint of the revolute joint type allowing only for rotation of the Dr_Sht component around the chosen axis,  Mi_Ni - joint of the revolute joint type allowing only for rotation of the Nut component around the chosen axis,  Screw_i - joint of the screw type creating a screw transmission between the components Nut and Screw,  Sfr_i - joint of the spherical type determining the possible motion between the components Screw and Frame,  Gear_i - joint of the gear type determining the method of motion transmission from the Dr_sht component on the Nut component,  Slider_i - joint of the slider type determining linear motion between trucks mounted on the Nut component and guides mounted on the Screw component"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000647_iciea.2015.7334244-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000647_iciea.2015.7334244-Figure7-1.png",
        "caption": "Fig. 7 Collision detection of cuboid-cuboid",
        "texts": [],
        "surrounding_texts": [
            "Therefore, object collision is simplified to geometry collision. In this paper, geometry is subdivided as the combination of point, line segment (denoted as line for convenience) and rectangle. Six kinds of geometry (point, line, rectangle, sphere, cylinder and cuboid) above are collectively called basic geometry, as shown in Fig. 2.\n0P\ncy\ncx\ncz\ncO\n2P\n1P 3P 4P\n5P\n6P 1s\n2s3s\n1P\n2P cO 1s\n2s\n0P\n1P\n2P r\nr\nFig. 2 Basic geometry\nFurthermore, we extracted concept of \u201ccenter distance\u201d \u2013 the distance between center of two geometry; and \u201cgeometry dimension\u201d \u2013 the maximum distance from center to surface of a geometry. The information of all basic geometry is shown in Table 1.\nIII. COLLISION DETECTION OF BASIC GEOMETRY After simplification of manipulator and environment, object collision can be classified into 4 kinds of collision \u2013 spherecylinder, sphere-cuboid, cylinder-cylinder and cuboid-cuboid collision. For convenience, cylinder-cuboid collision is seen as cuboid-cuboid collision. To skip collision detection between 2 geometry far away, we proposed a center distance criterion: if the center distance between 2 geometry is greater than the sum of their geometric dimension, there is no collision. Otherwise it need further detection. In addition, considering safety margin dS , the criterion is shown with following expression:\n( , ) . . i j i j Sif Dist G G D D d No collision else Need more detection > + +\nIn further detection, the 4 types of collision detection can be converted into 5 basic detections: point-line, point-cuboid, lineline, line-rectangle, line-cuboid. Their relation is given in Fig. 3.\nThe following is an introduction of the 4 types of collision detection. The algorithm returns 1 if any collision happens,; otherwise returns 0.\na) Apply center distance criterion on sphere-cylinder. The algorithm returns 0 if (1) meets, otherwise turns to b);\n2 2 0 1 2( ) = || || || || /4S C m C S SDist G ,G P P PP r r d> + + + (1)\nb) Compute the shortest distance dmin between point P0 to line P1P2.\nDo projection of point P0 to line P1P2. If 1 1 2/ [0,1]vPP PP\u03bb = \u2208 , then min 0|| ||vd P P= , otherwise\nmin 0 1 0 2 min{ || ||,|| ||}d P P P P= . c) Detect collsion of sphere-cylinder by dmin. The algorithm returns 0 if (2) meets; otherwise returns 1.\na) Apply center distance criterion on sphere-cuboid. The algorithm returns 0 if (3) meets, otherwise turns to b);\n2 2 2 0 1 2 3Dist( , ) || || || || || || || ||S C c S SG G P O r d= > + + + +s s s (3)\nb) Detect whether point P0 is outside of cuboid C. For any surface of cuboid with outward normal vector in\nand center point iP , if 0 0i iP P n\u22c5 < , then point P0 is outside, otherwise turns to c). The algorithm computes dmin of point P0 to the corresponding surface. The algorithm returns 0 if dmin > rS ; otherwise returns 1.\nc) Detect next surface.\n2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA) 935",
            "If all the 6 surfaces of cuboid fail to meet 0 0i iP P n\u22c5 < , then point P0 is inside of cuboid, which means 1 to return. 3) Cylinder \u2013 Cylinder\n1P\n2P\n1vP\nmind\n3P\n4P\n2vP\n1C\n2C\n2l 1l\n1mP 2mP\n1Cr 2Cr\n1mP\n4P\n2mP\n1vP\n2vP\nmind\n2P\n1P\n1 3 1 4 2 3 2 4 min{|| ||,|| ||,|| ||,|| ||}PP PP P P P P . c) Detect collsion of cylinder-cylinder by dmin. The algorithm returns 0 if (7) meets; otherwise returns 1.\na) Apply center distance criterion on cuboid-cuboid: The algorithm returns 0 if (8) meets, otherwise turns to b);\n2 2 2 1 2 1 2 3\n2 2 2 1 2 3\nDist( , ) || || || || || || || ||\n|| || || || || ||\nC C c c\nS\nG G O O '\n' ' ' d\n= > + +\n+ + + +\ns s s\ns s s (8)\nb) Detect whether vertex of C1 is inside of C2. The collision is the point-cuboid type. The algorithm\nreturns 1 if a vertex collides, otherwise detects next vertex. If all 8 vertexes have no collision, turns to c);\nc) Detect whether edge of C1 is in collision with C2. The collision is the line-cuboid type. The algorithm\nreturns 1 if an edge collides, otherwise detects next edge. If all 12 edges have no collision, returns 0.\nIV. COLLSION DETECION OF MANIPULATOR In the base of simplified geometric models and collision detection of basic geometry, we realized collision detection of manipulator. The algorithm flow chart is shown in Fig. 8.\nThe simplified model of a n-DOF manipulator is shown in Fig. 9.\nTo process collision detection of manipulator, we need the key information of basic geometry \u2013 key point, key vector and key length in an absolute reference system. In this paper, we took the base coordinate system of manipulator as the absolute reference system. Different geometry demands for different key information, as listed in Table 2. The expression of key point and key vector are same in Cartesian form. When joint angles are given, the relevant configuration, key point and key vector of manipulator can be solved by forward kinematics. The key length will not change without considering flexibility of manipulator.\n936 2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA)",
            "When a collision occurs, it's necessary to find out which component of links is in collision, so we have to number basic geometry of all links. Meanwhile, to improve efficiency of collision detection, we can exclude those links that will never collide according to construction of manipulator. For example, the components in one link or two adjacent links will never collide. Therefore, there's a collision detection list for every geometry, and we only need to process geometry in the list when traversing collision detection. The model system of a manipulator is shown in Fig. 10.\nV. SIMULATION AND EFFICIENCY OF THE ALGORITHM As mentioned in section II, we mainly focused on manipulators of cylinder and cuboid shape and the algorithm is in high precision under this condition. In order to verify the validity of the algorithm, we developed a simulation system based on OSG for manipulator, as shown in Fig. 11.\nThe D-H parameters of the manipulator are given in Table 3. The 8 links from the base to the end is defined L1 to L8, and L0 represents the base. Slide guide S and Flatbed F are seen as environment. Collision detection works rapidly by adopting the algorithm in this paper."
        ]
    },
    {
        "image_filename": "designv11_13_0003507_978-981-13-0305-0_5-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003507_978-981-13-0305-0_5-Figure6-1.png",
        "caption": "Fig. 6 Parameters related to cost",
        "texts": [],
        "surrounding_texts": [
            "The present stage describes the procedure for chemical post processing of the specimen. A chemical bath is prepared with acetone (Acetone 85% and distilled water 15%) as ABS plastic is soluble in acetone and also due to its low cost, low toxicity and good diffusion property. The specimen is now immersed for one hour. After removing it from ethylene, it is dried and weighed and its dimensions are measured. The surface roughness of the part is again measured to find out any difference in roughness."
        ]
    },
    {
        "image_filename": "designv11_13_0000469_ipec.2014.6869935-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000469_ipec.2014.6869935-Figure5-1.png",
        "caption": "Fig. 5. Simplified diagram of current spectrums caused by proposed method.",
        "texts": [
            " We discuss both c) and d) at the same time. At the case of the transition from.!high throughjiow tOthigh, a peak frequency of current spectrums is given in (9). 2 fpeak = 1 I -- + - !tow fhigh _ 2 !tow !t,;gh ,I;ow + .[high (9) At the case of the transition from.how thrOUgh.thigh tojimj', a peak frequency of current spectrums is given in (10). 2 fpeak = 1 I -- + - !tow fhigh _ 2 !tow !t,;gh ,I;ow + .[high (10) Therefore, current spectrums of the proposed method consist of three spectrums shown in Fig.5. That is a) the spectrums at the case of continuing jimj', b) the spectrums at the case of continuing .thigh, and c), d) the spectrums at the case of transition betweenhow and .!high. Each spectral peaks of the three spectrums are given in (II), (12), and (13). fhi\"h Phi (1- Plh) Snectral peak of I\" = -'_ . c y . .11\",< I\" I\" P I\" P J 10\\\\' J 10lt' ih + J JlIgh hi !t P(I-P) Snectral peak of 1\". = low Ih hi C y . .I h/[(h I\" I\" P, I\" P, J high J low Ih + J high h1 (11) (12) (13) Where, C is common value that is changed according to modulation factor,Ji,igh,jiOlv, and so on"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002624_s00332-016-9299-4-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002624_s00332-016-9299-4-Figure7-1.png",
        "caption": "Fig. 7 Direction and intensity (proportional to the length of the arrows, arbitrary units) along the filament midline of the three perturbation modes entering (58) for n = 2: (a) in-plane perturbation; (b) twisting perturbation; (c) out-of-plane perturbation",
        "texts": [
            " With the above definitions, the specialization of the second variation (55) for this solution becomes \u03b42 [\u03be\u0302 , d\u0302](v,w, h) = \u222b 2\u03c0 0 \u222b 1 0 \u03c12w2 \u03c1 + w2 \u03b8 \u03c1 d\u03c1d\u03b8 + \u222b 2\u03c0 0 (v2 \u2212 v2\u03b8 )d\u03b8 + \u222b 2\u03c0 0 \u03b11(v\u03b8\u03b8 + v)2d\u03b8 \u2212 \u222b 2\u03c0 0 (1 + \u03b11(1 \u2212 \u03bc\u03042))w 2 \u03b8d\u03b8 + \u222b 2\u03c0 0 (\u03b12 \u2212 \u03b11(1 \u2212 \u03bc\u03042))h 2d\u03b8 + \u222b 2\u03c0 0 (2\u03b12 \u2212 2\u03b11(1 \u2212 \u03bc\u03042))hw\u03b8\u03b8d\u03b8 + \u222b 2\u03c0 0 \u03b12w 2 \u03b8\u03b8d\u03b8 + \u222b 2\u03c0 0 \u03b13(h\u03b8 \u2212 w\u03b8) 2d\u03b8. (57) Due to the invariance of our problem with respect to global translations and rotations, we can expand the test fields on a complete orthonormal basis in their function space as follows: v(1, \u03b8) = 1\u221a \u03c0 \u221e\u2211 n=2 An cos n\u03b8, h(\u03b8) = 1\u221a 2\u03c0 B0 + 1\u221a \u03c0 \u221e\u2211 n=1 Bn cos n\u03b8, w(\u03c1, \u03b8) = 1\u221a \u03c0 \u221e\u2211 n=2 cn(\u03c1) cos n\u03b8. (58) We also set Cn := cn(1) to simplify various formulae. We illustrate in Fig. 7 the three perturbation modes entering (58) corresponding to n = 2: As we show below, these are the perturbations associated with the first buckling mode of the system. On substituting (58) into (57) and introducing the curvature mismatch \u03b6 defined in (52), the stability condition (41) reduces to the requirement that the inequality \u03b42 [\u03be\u0302 , d\u0302](An, Bn,Cn, cn) = \u221e\u2211 n=2 (1 \u2212 n2 + \u03b11(1 \u2212 n2)2)A2 n + \u221e\u2211 n=0 (\u03b12 + \u03b13n 2 \u2212 \u03b11\u03b6 )B2 n + \u221e\u2211 n=2 n2(\u03b13 \u2212 1 \u2212 \u03b11\u03b6 + \u03b12n 2)C2 n \u22122 \u221e\u2211 n=2 n2(\u03b13 + \u03b12 \u2212 \u03b11\u03b6 )BnCn + \u221e\u2211 n=2 \u222b 1 0 ( \u03c1c2n,\u03c1 + n2 \u03c1 c2n ) d\u03c1 \u2265 0, (59) holds for all choices of An , Bn , Cn , and cn ",
            " Restricting attention to the case n = 2 and dividing (65)1 and (65)2 by \u03b11 > 0 and (65)3 by \u03b12 1, after simple algebra we obtain the desired necessary and sufficient conditions for the stability of the flat circular disk, namely \u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 \u03bd \u2264 3, \u03bd \u2264 2(4\u03b22 \u2212 \u03b6 + \u03b23), \u03bd \u2264 6\u03b6(\u03b22 \u2212 \u03b6 + \u03b23) + 18\u03b22\u03b23 \u03b22 \u2212 \u03b6 + 4\u03b23 , (66) where we have introduced, along with the dimensionless parameters \u03b6 and \u03b22 defined in (52), the further dimensionless parameters \u03bd := 1 \u03b11 and \u03b23 := \u03b13 \u03b11 , (67) which, respectively, measure the relative importance of the force generated by surface tension and the largest bending rigidity and the ratio between the twisting rigidity and the largest bending rigidity. We observe that in the limiting case \u03b6 = 1 (vanishing intrinsic curvature) it is essential to confine attention to situations where the rigidities \u03b11 and \u03b12 are equal, so that \u03b22 = 1. In this case, the relevant stability threshold is given by (66)1 and the first instability is generated by purely in-plane perturbations (Fig. 7a). It is interesting to study the critical value \u03bdcr of the dimensionless effective surface tension at which the instability sets in. This quantity is defined, as a function of the rigidity ratios \u03b22 and \u03b23 and of the curvature mismatch \u03b6 , by \u03bdcr(\u03b22, \u03b23, \u03b6 ) := min { 3, 6\u03b6(\u03b22 \u2212 \u03b6 + \u03b23) + 18\u03b22\u03b23 \u03b22 \u2212 \u03b6 + 4\u03b23 } . (68) From (68), it is evident that the parameter region in which the first buckling mode is a purely in-plane deformation is characterized by the condition \u03bdcr = 3, while the first instability is given by a buckling mode with out-of-plane (Fig. 7c) and twisting (Fig. 7b) components when \u03bdcr < 3. In Fig. 8, we plot \u03bdcr as a function of \u03b22 and \u03b23 for different values of \u03b6 . It can be seen that the region where the first buckling mode is an out-of-plane plus twisting one (\u03bdcr < 3) becomes smaller and smaller upon increasing \u03b6 and disappears for \u03b6 \u2265 1/2. We note that our analysis encompasses also the case \u03b6 < 0 and we can infer from (68) that, under this condition, the solution becomes unstable even in the absence of the soap film for sufficiently small values of \u03b22 and \u03b23"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003423_physrevfluids.3.043101-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003423_physrevfluids.3.043101-Figure8-1.png",
        "caption": "FIG. 8. Changes of the first-order swimming velocity U1/U0 in (a) the \u03b22-\u03b23 plane and (b) the \u03b22-\u03b24 plane, assuming Pe = 400. Here \u03b21 is calculated from 2 3 \u03b22 1 + \u2211 i =1 \u03b22 i = 1. It is assumed that \u03be = 1.",
        "texts": [
            " Nonetheless, the changes in U1/U0 for the combination of the first and third modes are smaller than what has been observed for the combination of the first and second surface stroke modes. Also U1/U0 curves of higher surface stroke modes, |\u03b24|/\u03b21 in Fig. 7(f), show minor sensitives to the sign of the surface stroke mode. At high P\u00e9clet number of Pe = 500, U1/U0 shows great sensitivity to the magnitude and sign of \u03b2i/\u03b21 for all test cases [Figs. 7(g)\u20137(i)]. Moreover, the smallest U1/U0 is observed in all cases at |\u03b2i |/\u03b21 3 for either positive or negative higher stroke modes. Finally, to see the effects of surface motion with more than two stroke modes, in Fig. 8 the changes of U1/U0 within the \u03b22-\u03b23, \u03b22-\u03b24, \u03b22-\u03b25, and \u03b23-\u03b25 planes are plotted. In each case, the boundary of search space and the value of \u03b21 are from the constraint of the invariant consumed energy [Eq. (16)]. Moreover, here we choose \u03be = 1 and Pe = 400. The (0,0) node in each plane corresponds to the pure treadmill mode, which previously had been shown to be both the optimal swimming and optimal feeding scenario [13]. It is observed that the maximum U1/U0 does not correspond to the pure treadmill surface stroke and happens for a finite amplitude of higher modes",
            "), the maximum of U1/U0 corresponds to the case in which both modes have the same sign, but if one of the higher modes is odd and the other is even, the maximum of U1/U0 occurs when the modes have opposite signs. The maximum value of U1/U0 for these cases is larger and reaches 0.044 compared to U1/U0 = 0.035 043101-14 in the case with pure treadmill motion. This is evidence for more optimal swimming strokes than just pure treadmill when a nonuniform viscosity environment is considered. The streamlines and concentration contours of cases corresponding to the maximum U1/U0 observed in Fig. 8 are shown in Fig. 9. The combination of the surface stroke modes in these cases makes the constructive cm (cm < 0) stronger at the front and back of the swimmer, mainly by increasing the radial gradient of c in those spots. At the same time, the optimal combination of the modes marginally decreases the strength and extent of the destructive cm (cm > 0). The unique feature observed for all optimal combinations of higher modes is that the boundary layer stays connected and higher surface stroke modes only marginally modulate the boundary layer of the pure treadmill stroke without inducing any new stagnation point along the surface of the swimmer. The minimum U1/U0 of the cases shown in Fig. 8 is less than \u22120.081 with the minimum observed value of U1/U0 = \u22120.097 corresponding to surface motion of \u03b21 = 0.454, \u03b22 = 0.869, and \u03b23 = \u22120.328. It is found that the regions with a negative U1/U0 value are narrow in the radial direction and elongated in the angular direction. The contour plots of c and cm for the cases with minimum U1/U0 are shown in Fig. 10, where the boundary layer modifies substantially with new stagnation points appearing along the surface of 043101-15 the swimmer. It reduces the size and strength of the regions with negative cm and enlarges regions with positive cm toward the front of the swimmer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003363_1.g002683-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003363_1.g002683-Figure2-1.png",
        "caption": "Fig. 2 Flap and skewed lag-pitch hinges are located coincident at radius eR.",
        "texts": [
            " At the same time, the blade tip on the left would be forced to flap upward twice per revolution as that blade obtains its maximum lead and lag angles. The resulting flapmotions for the two blades could not bematched, and these undesirable higher harmonics in the flapping response would contribute to large bending moments at the blade roots and unwanted airframe vibrations. The addition of a flap hinge relieves this kinematic constraint and allows the smooth sinusoidal flapping motion and conventional tip path plane response evident in the experiments. The simplified kinematics are depicted in Fig. 2 with respect to a rotating hub-fixed coordinate system with unit vectors {x\u0302, y\u0302, z\u0302}. The kinematics and coordinate conventions for the positive lag-pitch coupling blade are shown on the right side of the figure. The hub rotates about the z\u0302 axis by angle \u03c8 with respect to an inertial frame. The inboard flap hinge axis is fixed in the hub body and joins the cross body. The flap hinge rotates by an angle \u03be1 about an axis pointed in D ow nl oa de d by T U FT S U N IV E R SI T Y o n Fe br ua ry 1 5, 2 01 8 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .G 00 26 83 the\u2212y\u0302 direction. The lag hinge axis of rotation is inclined by an angle \u03b4 from vertical to point in the sin \u03b4 x\u0302 \u2212 cos \u03b4 z\u0302 direction, and the hinge rotates by an angle \u03be2. The flap hinge and lag hinge are collocated at radius eR for blade tip radius R and eccentricity 0 < e < 1. We would like to make a precise analogy between the actual kinematics of Fig. 2 and the conventional parameterization of blade motions in terms of orthogonal lag and flap axes. To do this, we consider small deflections of the blade about its physical hinges. The composite rotation about first and second axes is conveniently described by exponential coordinates (or an axis and angle representation) when the rotations are infinitesimal. A finite rotation by angle \u03be1 about an axis with unit vector \u03c91 is described by the rotation matrix exp \u03c9\u03021\u03be1 , where \u03c9\u03021 is the skew symmetric matrix defined such that \u03c91 \u00d7 b \u03c9\u03021b for all b. For the case of an infinitesimal rotation size d\u03be1, then to a first-order approximation exp \u03c9\u03021\u03be1 I \u03c9\u03021d\u03be1. It follows that the composite rotation about axis \u03c91 by angle d\u03be1 and then about axis \u03c92 by angle d\u03be2 is exp \u03c9\u03022d\u03be2 exp \u03c9\u03021d\u03be1 I \u03c91d\u03be1 \u03c92d\u03be2 \u2227 to a first-order approximation. A physical interpretation of this result is simply that infinitesimal rotations commute, or that velocity vectors add. The exponential coordinates for the composite rotation dictated by the design geometry of Fig. 2 are expressed by 2 4 0 \u22121 0 3 5d\u03be1 2 4 sin \u03b4 0 \u2212 cos \u03b4 3 5d\u03be2 (1) For analysis, we re-parameterize the motion in terms of the canonical flap angle \u03b2 about an axis in the\u2212y\u0302 direction and lag angle \u03b6 about the \u2212z\u0302 direction, with both axes fixed in the hub frame. This arrangement is shown in Fig. 3. We separately impose a geometric lag-pitch coupling coefficient \u0394\u03b8\u2215\u0394\u03b6 tan \u03b4 and the resulting exponential coordinates for the composite rotation are 2 4 0 \u22121 0 3 5d\u03b2 2 4\u0394\u03b8\u2215\u0394\u03b6 0 \u22121 3 5d\u03b6 (2) The reparameterized expression encodes identical kinematics constraints as the original",
            " These effects are lumped into equivalent nondimensional linear damping coefficients c\u03b2 and c\u03b6 for flap and lag in the analysis. Instead of fitting these parameters from data, reasonable estimates are derived by an energy argument that highlights some expected scaling relations for these coefficients. The pin joints are principally loaded by the outward centrifugal force F of the spinning blade, which can be computed by integrating \u03a92R\u03bedm over the mass of the blade. F 1 2 mR 1 e \u03a92 3 2 I\u03b2\u03a92 1 R 1 e 1 \u2212 e 2 (26) The physical flap hinge of Fig. 2 is modeled as a plain journal bearing or short shoe brake, for which the friction torque \u03c4\u03be1 depends Table 1 Propeller properties Parameter Symbol Value Tip radius R 159 mm Number blades Nb 2 Hinge eccentricity e 0.076 Washer disk radius RD 1.98 mm Hinge pin radius RP 0.52 mm Blade mass m 5.40 g Hub rotational inertia \u2014 \u2014 5.1 \u00d7 10\u22127 kg \u22c5m2 Blade chord c 19.3 mm Blade pitch \u03b80 9\u00b0 Section drag coef cd0 0.06 Section lift curve slope a 0.1\u2215 deg Friction coef steel\u2013plastic \u03bc1 0.20 Friction coef plastic\u2013plastic \u03bc2 0",
            " In addition, a large axial load Fj sin \u03b4 j is carried by two plastic washers of radius RD, which slide against each other with material coefficient of friction \u03bc2. This contributes a second term to the friction torque about the skew axis, which is modeled as the torque of a uniform pressure contact disk brake [15]. Once again a geometric parameter GD for these disks is introduced such thatRD GDR. The friction workW\u03be2 of torque \u03c4\u03be2 integrated over one cycle of lag amplitude A\u03b6 is computed, recognizing from the geometry of Fig. 2 that the skew hinge axis rotates with an amplitude A\u03b6\u2215 cos \u03b4 . \u03c4\u03be2 \u03bc1GPRF cos \u03b4 2 3 \u03bc2GDRFj sin \u03b4 j (32) W\u03be2 6I\u03b2\u03a92A\u03b6\u03bc1GP 1 e 1 \u2212 e 2 4I\u03b2\u03a92A\u03b6\u03bc2GD 1 e 1 \u2212 e 2 j tan \u03b4 j (33) As before, setting this friction work expression equal to a damping work expression allows an equivalent nondimensional damping coefficient c\u03b6 to be defined for the conventional lag coordinate in the dynamics. c\u03b6 6 \u03c0 1 A\u03b6 \u03bc1GP 2 3 \u03bc2GDj tan \u03b4 j 1 e 1 \u2212 e 2 (34) The final contribution to the overall dynamics in Eq. (6), \u03c4hinge, is nowwritten in terms of these damping coefficients and the coordinate velocities _x \u03c9; _\u03b6; _\u03b2 ",
            " Synchronization of these sources is achieved by having the motor controller emit a digital index signal read by the load cell DAQaswell as a visible indicator for the high-speed video. The equation ofmotion developed for the half-propeller in Eq. (50) takes advantage of approximate symmetry to describe the dynamics of a single blade instead of explicitly modeling two blades. To practice cyclic control, one of the blades is mounted with a positive lag-pitch coefficient\u0394\u03b8\u2215\u0394\u03b6 and the other is mounted with a negative lag-pitch coefficient, as shown in Fig. 2. Model predictions for both the positive and negative cases are plotted against the measured data\u2014the difference is only notable in the case of flap, where the purpose of the cyclic system is to ensure that the positive and negative blades remain 180\u00b0 out of phase with each other. The model state space could be extended to explicitly encompass the full systemwith two independent dissimilar blades, but the simplified model used here exposes the fundamental physics being exploited and makes satisfactory numerical predictions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000152_boca-2014-0001-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000152_boca-2014-0001-Figure3-1.png",
        "caption": "Figure 3: Effect of pH on activity of free (u) and immobilized lipase (n). Experiments were performed in triplicate. The error bars indicate the standard deviation (3%).",
        "texts": [
            " This value was in agreement with enzyme immobilization efficiency, calculated to be 85%. The protein loading amount was ~ 3.4 mg/g carrier (3468 U/g carrier). Unauthenticated Download Date | 1/12/15 9:57 PM Unauthenticated Download Date | 1/12/15 9:57 PM The catalytic properties of the free and immobilized lipase were studied under various pH conditions and temperatures. The optimum pH for the enzymatic activity of free lipase was in the range 6.0 \u2013 6.5 and the maximum activity for the immobilized lipase was found to occur at a pH of 6.5 (Fig. 3). As a result of the immobilization process, the catalytic activity of the immobilized enzyme was maintained around 90%, while, for the free lipase, a significant activity decrease to 40% was detected. The temperature-activity profile of the free and the immobilized lipase was investigated to determine the matrix-protecting effect on the immobilized enzyme with respect to thermal denaturation. Maximum activities were observed at 40\u00a0\u00b0C for the free and immobilized enzyme, as reported in Fig. 4. The temperature profile indicated a higher activity for the immobilized enzyme (55 \u2013 60 \u00b0C) compared to that of the free form"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003705_ecce.2018.8558256-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003705_ecce.2018.8558256-Figure1-1.png",
        "caption": "Fig. 1. Sketch of the ideal three-phase equivalent rotor winding.",
        "texts": [
            " In order to prove the accuracy of the proposed direct analysis strategy, a comparison with measurements is reported, considering a standard, double cage, grid fed IM, at different loads. Measurements have been done according to the current standard for AC machines efficiency computation [14]. In a squirrel cage IM, spatial distributions of induced voltage and current in the rotor bars exhibit almost a sinusoidal waveform. Then, an interesting trick is considering a threephase rotor equivalent winding sinusoidally distributed in the rotor slots. Fig. 1 shows a sketch of such a three-phase rotor winding. The shading indicates the theoretical number of conductors belonging to each phase within the rotor slots. As a consequence, imposing a three-phase current in the rotor winding, this assumption yields a sinusoidally distributed current in the rotor bars. The rotor turns per phase are computed from the magneto- motive force (MMF) balancing as: Nr kwr = Ns kws (1) where kws and kwr are the stator and rotor winding factors and Ns is the number of turns per phase of the stator winding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000489_physrevlett.108.074301-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000489_physrevlett.108.074301-Figure1-1.png",
        "caption": "FIG. 1. A compressing device (a) involves cylindrical plates distant from a controlled gap Y and a sheet clamped on their curved sides in between. (b) By iterative buckling, gap reduction yields the generation of folds of ever smaller size , whose axis is bent by the cylinders. This results in two principal curvatures and thus in Gaussian curvature and in-plane stretching.",
        "texts": [
            " We show in this Letter that it thus misses some interesting features displayed on compaction routes involving significantly different principal curvatures. In particular, the sheet may then undergo a spontaneous reverse uncrumpling at large anisotropic compaction. In the experiment, a thin polycarbonate planar sheet (length l \u00bc 155 mm, width L \u00bc 190 mm, thickness h ranging from 0.05 to 0.5 mm) is clamped along its largest side l onto the arches of a transparent upper cylinder whose curvature radius R \u00bc 50 cm is large (Fig. 1). As the distance between the clamping arches, X \u00bc 180 mm, is smaller than the sheet width L, the sheet is already buckled before compression (Fig. 2). A bottom cylinder, parallel to the upper one, is then moved up so as to impose a gap Y between them. Further evolution of the gap Y is then set and controlled by stepper motors to an accuracy of a tenth of microns [16,17], thus allowing a detailed study of the compaction. Whereas one of the principal curvature of the sheet c1 is induced by iterated buckling, a second one c2 is imposed by the cylinders curvature (Fig. 1). Both yield a PRL 108, 074301 (2012) P HY S I CA L R EV I EW LE T T E R S week ending 17 FEBRUARY 2012 0031-9007=12=108(7)=074301(4) 074301-1 2012 American Physical Society nonzero Gaussian curvature G \u00bc c1c2 and stretching, in any states. In view of the large curvature radius (R Y), it is instructive to consider the specific case of compaction between parallel plates [16\u201318], i.e., R \u00bc 1. An isometric solution with the sheet shape invariant along the clamping direction is then allowed. When reducing the gap Y, it gives rise, by iterated buckling, to successive generations of regular and identical folds. In particular, a n-fold solution can be formally constructed by repeating n times a single fold of height Y but scaled down by a factor 1=n. It is then made of n folds of reduced height Y=n and width \u00bc X=n, meaning that n Y 1 1. In comparison, using a large but finite R here means weakly bending the axes of these folds [Fig. 1(b)]. However, as R Y, the above geometric scaling laws are expected to hold, at least on smooth uncrumpled states. As the gap Y is reduced, buckling makes the sheet actually display an increasing number n Y 1 of folds whose right and left sides appear, respectively, in red or blue in Fig. 3. Folds first involve a contact surface with the cylindrical plates ending by curved ridges [Figs. 3(a) and 3(b)], similar to those found in the uncompressed state of Fig. 2. These ridges condense the sheet\u2019s stretching on a curved line, similarly to a buckled ping-pong ball [19,20]",
            " This yields 2 for in-plane stresses, 4 EG from (1) and finally, 2EG. The stretching energy density then scales like es h 2=E Eh 4G2, so that the energy density ratio writes es=eb 4 with \u00bc \u00f0G=hC\u00de1=2: (2) This scaling estimate on regular states applies to all situations of compaction, isotropic or anisotropic. In line with the analysis of isotropic compression, one expects singular states for 1, regular states for & 1 and a transition between them for of order 1. Interestingly, in the present compaction by cylinders (Fig. 1), the energy densities of the smooth state can also FIG. 2. Clamped sheet prior compression showing two ridges at the contact lines between the sheet and the cylinder. PRL 108, 074301 (2012) P HY S I CA L R EV I EW LE T T E R S week ending 17 FEBRUARY 2012 074301-2 be derived exactly, assuming invariance along the cylinder orthoradial direction z of both the sheet shape 0 and the sheet stresses ( i;j). Taking into account the cylinder shape y \u00bc z2=2R, the sheet surface thus writes \u00f0x; z\u00de \u00bc 0\u00f0x\u00de \u00fe z2=2R, where 0 is a -periodic function with zero mean value, denoting the fold width"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001673_s11837-018-3242-0-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001673_s11837-018-3242-0-Figure7-1.png",
        "caption": "Fig. 7. Magnetic flux density distribution of a 20-W SPM BLDCmotor with (a) unaligned anisotropic bonded magnets with 65 vol.% loading of NdFe-B in nylon-12 and (b) anisotropic bonded magnets aligned at 0.25 T with 40 vol.% loading of Nd-Fe-B in nylon-12.",
        "texts": [
            "25 T predicted achievement of torque and power above the rated values, implying that the loading fraction of Nd-Fe-B in nylon-12 could be reduced, or that a smaller design might be practical. The properties of the anisotropic bonded magnet aligned at 0.25 T were scaled for lower loading fractions, and FEA was used to predict the minimum volume fraction at which the rated power could be achieved. The FEA results indicated that 40 vol.% loading of Nd-Fe-B in nylon-12 would be sufficient to achieve the rated power (Fig. 6), if aligned at 0.25 T. Figure 7 compares the magnetic flux density of the 20-W SPM BLDC motor when using magnets under isotropic conditions (unaligned) with 65 vol.% loading and under aligned anisotropic conditions (0.25 T) with 40 vol.% loading. The predicted remanence of the aligned magnet (0.29 T) with reduced volume fraction of 40% Nd-Fe-B is comparable to that of the unaligned magnet (0.30 T) with volume fraction of 65% Nd-Fe-B, accounting for the similarity in the magnetic flux density distribution. We were unable to validate the magnetic properties of the 40 vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000174_j.snb.2013.10.089-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000174_j.snb.2013.10.089-Figure3-1.png",
        "caption": "Fig. 3. (A) Scheme of the humid electronic n",
        "texts": [
            " erve gases and their mimics were degraded by water hydrolysis in arious steps until phosphoric acid or alkylphosphonic acids were iven [47]. As shown in Fig. 2, the first step was fast and produced In our particular case these reactions were expected to improve given the use of a weak base (i.e., borate) as a background salt. Furthermore, as detection was carried out in an aqueous environment, therefore, the gases that were not miscible in water and humidity will not interfere with the measurements. The \u201chumid electronic nose\u201d system (Fig. 3a) consists of one polypropylene piece formed by two cylindrical parts. The upper cylinder acts as top cover of the cell chamber and contains three inlets; two of them are used as an inlet and outlet for air, argon and vacuum (item 1 in Fig. 3a), and the third one is for the introduction of liquids and solids (item 2 in Fig. 3a). All the electronic connections between the equipment and electrodes are in this part, outside the measurement chamber. The lower cylinder was used as a support for the voltammetric measurement elements. It has three inlets: two are used for two different sets of working electrodes (item 3 in Fig. 3a) and the third one is used for the salt bridge connection with the reference electrode (item 4 in Fig. 3a). At the bottom there is a plane steel plate screwed as a counter-electrode (item 5 in Fig. 3a). The fabric of the membrane (item 6 in Fig. 3a) was made of nylon with a fibre thickness of 350 microns, separated by about 150 microns. It is fixed with a second polypropylene piece (item 7 in Fig. 3a) that fits the inferior cylinder holding, and tenses the membrane and secures the connection with all the electrodes and the salt bridge. Each time a new measurement was taken the membrane was rinsed with distilled water, dried and then moistened again with the salt solution. The measurement chamber (Fig. 3b) consisted in an 800-mL glass cell and the \u201chumid electronic nose\u201d system acted as a tap cover to seal the system. The cell was designed so that no gas leaks take place and it was thermostated with a recirculating temperature controller. The samples inlet was closed with a septum cap and liquid samples were introduced by pricking with a needle. Gas samples were introduced by the gas inlet. The \u201chumid electronic nose\u201d use an array of eight working electrodes (Ir, Rh, Pt, Au, Ag, Co, Cu Ni) with a purity of 99"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002005_toh.2014.2330300-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002005_toh.2014.2330300-Figure2-1.png",
        "caption": "Fig. 2. Hardware for multi-point augmentation.",
        "texts": [
            " The forces at the fingers are given as fH; \u00f0t\u00de \u00bc fR; \u00f0t\u00de \u00fe fT; \u00f0t\u00de: (1) The main task is to properly estimate and render the additional tumor forces, fT; , to create a sound illusion. As hardware setup we employ the same as the one used for two-point stiffness modulation described in [25]. In brief, the hardware consists of two impedance-type haptic devices (PHANToM premium 1.5; Geomagic) and customdesigned probing rods attached to each PHANToM. The tools are instrumented with 3D force/torque sensors (Nano17; ATI Industrial Automation, Inc.) as shown in Fig. 2. The diameter of the tip is 15 mm. Before describing the approach in detail, a few assumptions and simplifications should be indicated. The rendering method is in general independent of the type of manipulation. In our current testbed, the interaction is carried out via sensor-equipped probing rods. But with appropriate sensing and actuation hardware, the same approach could also be used for direct finger interaction. A further point to note is that we employ a homogeneous silicone hemisphere (104 mm in diameter) as base object"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001136_s00170-017-0625-2-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001136_s00170-017-0625-2-Figure6-1.png",
        "caption": "Fig. 6 Sections for metallographic observation in a thin wall and b ladder block. Section X is normal to the scanning direction, while section Y is parallel to the scanning direction",
        "texts": [
            "5, and the beads in a particular layer are always deposited exactly over the corresponding beads of the previous layer, as shown in Fig. 4b. Herein, the \u201coverlap ratio\u201d is recorded as OVr, and calculated using Eq. 1: OVr \u00bc OL=W \u00f01\u00de where OL is overlap width and W is single bead width, as shown in Fig. 4b. Furthermore, for recording the temperature history of the molten pool in the deposition process, infrared thermography was applied. The corresponding method is illustrated in Fig. 5. The FLIRX6530sc was the infrared thermo-graph used in this study. Figure 6 shows the different sections for metallographic observation, which are represented as sections X and Y. Section X is normal to the scanning direction, while section Y is parallel to the scanning direction. All the cross-sections were made by wire cutting. The metallographic samples were mounted, ground, and polished. The sample sections were electrically etched by 10% oxalic acid and 3\u20134 V electrolysis voltage for about 30 s. Microstructures were observed by an ordinary optical microscope and a laser confocal microscope"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003508_0037549718784186-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003508_0037549718784186-Figure1-1.png",
        "caption": "Figure 1. Schematic diagram of the MM. MM: mobile manipulator.",
        "texts": [
            " In section 3, the bond graph modeling of the MM is explained. Section 4 discusses the control strategies for the MM. Simulation results and conclusions are presented in section 5 and 6 respectively. The physical model of the MM consists of two major parts viz. an MB and a planar MA. The MB has three omni wheels, and each wheel is independently actuated by a separate DC motor; being equipped with three motors, the MB is a holonomic device. The horizontal planar MA has four degrees of freedom. The schematic diagram of the MM is shown in Figure 1, illustrating the different reference frames in the XY plane. The motion of the MB is considered with reference to the world coordinate frame {W}, the MB coordinate frame {G} is placed at the center of mass (CM) of the MB, and {E} is the frame of the tip of the MA. P is the point at which the MA is attached to the MB and {P} is the angular absolute frame parallel to {W}. The three-wheeled MB has been modeled by several researchers in the past.9\u201311 Figure 2 shows the velocities of the wheels and the CM of the MB",
            " The kinematic equations relating the linear velocities of wheels with the CM velocities of the MB as shown in Figure 2 can be written as: _c1 _c2 _c3 2 4 3 5= 1 r sinf cosf L sinf1 cosf1 L sinf2 cosf2 L 2 4 3 5 Vx,w Vy,w _f 2 4 3 5, where, f1 = p 3 f and f2 = p 3 +f \u00f01\u00de The planar MA is appended to the MB at point P which is at a distance LP from the CM of the base. L1, L2, L3 and L4 are lengths of link-1, 2, 3 and 4 respectively, and u1, u2, u3 and u4 are corresponding joint angles, as shown in Figure 1. Since the MA is attached to the MB with an angular displacement f, the coordinates of each link are expressed in the frame {P} for development of the bond graph model of the MA. The coordinates of the tip of the link-1 in the frame {P} are as follows: X1,P = L1 cos (f+ u1) \u00f02\u00de Y1,P = L1 sin (f+ u1) \u00f03\u00de The velocities of the tip of the link-1of MA in X and Y direction with respect to the CM of the MB are derived as follows: _X1,P = L1 sin (f+ u1)( _f+ _u1) \u00f04\u00de _Y1,P = L1 cos (f+ u1)( _f+ _u1) \u00f05\u00de In a similar fashion, the velocities of the tips of link-2, link-3, and link-4 respectively, could be written as follows: _X2,P = _X1,P L2 sin (f+ u12)( _f+ _u12) \u00f06\u00de _Y2,P = _Y1,P + L2 cos (f+ u12)( _f+ _u12) \u00f07\u00de _X3,P = _X2,P L3 sin (f+ u123)( _f+ _u123) \u00f08\u00de _Y3,P = _Y2,P + L3 cos (f+ u123)( _f+ _u123) \u00f09\u00de _X4,P = _X3,P L4 sin (f+ u1234)( _f+ _u1234) \u00f010\u00de _Y4,P = _Y3,P + L4 cos (f+ u1234)( _f+ _u1234) \u00f011\u00de where, u12 = u1 + u2, u123= u12 + u3, and u1234 = u123 + u4, _u12 = _u1 + _u2, _u123= _u12 + _u3, and _u1234 = _u123 + _u4 Bond graph is a technique for modeling of the dynamics of physical systems"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003387_s10586-018-2292-y-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003387_s10586-018-2292-y-Figure1-1.png",
        "caption": "Fig. 1 Mechanical control leveling system schematic diagram",
        "texts": [
            " In this paper, the whole vehicle active suspension system is taken as the object, and the whole vehicle active suspension system model with full consideration of the dynamic characteristics of hydraulic device is established. Aiming at the nonlinear term of hydraulic device in the model, a control strategy combining sliding mode control algorithm with Backstepping method is put forward. Finally, the simulation results are given and analyzed. Mechanical control leveling system has relatively simple structure, which mainly relies on an external hydraulic source to control the amount of oil into the hydraulic cylinder through the open and close of the mechanical control valve group. As shown in Fig. 1, the height of the suspension is controlled by the relative displacement between the axle and the body. When the valve set position is in line with the design value, the hydraulic cylinder is connected with the hydraulic pump or fuel tank, so that the height of the body is always maintained at the design height. As shown in Fig. 1, the connecting rod is used to connect the three- position three-way valve with the axle, and the valve group is fixed on the body. When the body and the vehicle axle have vertical relative movement, the valve group will open or close, thus controlling the hydraulic cylinder oil filling or drain oil [11]. For example, when the vehicle load increases, the body is depressed, and the relative displacement between the valve group and the axle decreases. At this point, the right side of the valve group is connected to the high-pressure source on the left"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003941_acs.analchem.9b01912-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003941_acs.analchem.9b01912-Figure1-1.png",
        "caption": "Figure 1. Description of the setup combining a CAP source and the shielded electrochemical cell for in situ measurements. (A) Profile schematic view (not at scale) of the experimental setup including the plasma source equipped with a device providing a surrounding gaseous environment,12 the cuvette containing the solution to be exposed (6 mL of PBS), and the two shielded containers for electrochemical measurements with two counter electrodes (CE 1 and CE 2) in container 1 and two Pt working electrodes (WE 1 and WE 2) together with the reference electrode (Ag/ AgCl, noted REF) in container 2. Characteristic distances are a = 20 mm, b = 7 \u00b1 1 mm, c = 12 \u00b1 1 mm, and d = \u223c20 mm. (B) Top view showing the position of the electrodes in the containers (CE 1 and CE 2 in container 1; WE 1, WE 2, and REF in container 2) and in the quartz cuvette with respect to the plasma impact point and the magnetic stirrer (not at scale). (C) Picture of the experimental setup under operation (see the plasma wave generated in the tube, which propagates and reaches the solution).",
        "texts": [
            " PBS was from PAN BIOTECH (10 mM, pH 7.4, KCl 0.003 mol\u00b7L\u22121; KH2PO4 0.002 mol\u00b7L\u22121; NaCl 0.137 mol\u00b7L\u2212\u20111; Na2HPO4 0.01 mol\u00b7L\u22121, without Mg2+/Ca2+). Solutions of H2PtCl6, lead acetate, Fc(MeOH)2, H2O2, and NaNO2 were from Sigma-Aldrich (France). Plasma Treatment of Physiological Solutions. Six milliliter PBS solutions were exposed to plasmas in a homemade quartz cuvette (external diameter 44 mm, inner diameter 41 mm, depth 40 mm). A distance of 20 mm was fixed between the end tube of the plasma device and the solution (see Figure 1A characteristic distance \u201ca\u201d in Figure 1A). Moreover, the cuvette was placed over a metallic support connected to the ground in order to ensure a constant potential to the sample. Plasma Setup. The plasma setup used in these experiments has been described previously.12 A surrounding gas device allows us to control the gaseous environment of the gas phase (plasma). We used helium as the working gas (1.67 slm rate) and the environment gas was composed of 100% N2 (0.03 slm rate; Linde, 99.9995%), for a total gas flow of 1.7 slm. The helium flow rate was adjusted with adequate mass flow controllers (EL-FLOW, Bronkhorst High-Tech) connected to a flow-bus",
            " All electrochemical measurements were carried out using a bipotentiostat (BioLogic, VSP-300, EC-Lab software) equipped with low current modules kept in a homemade Faraday cage. In order to detect RONS generated in PBS during plasma exposure, cyclic voltammograms were achieved between \u22120.1 and +0.9 V vs Ag/AgCl, at 20 mV\u00b7s\u22121 scan rate.12 Chronoamperometry was performed at either +0.3 or +0.85 V vs Ag/AgCl. Moreover, magnetic stirring was necessary to increase convection of exposed solution from the plasma impact point in the PBS to the electrode surface (Figure 1). \u25a0 RESULTS AND DISCUSSION Development of the Experimental Setup Based on a Shielded Electrochemical Cell. The final goal of this work is to perform sensitive electrochemical measurements in an DOI: 10.1021/acs.analchem.9b01912 Anal. Chem. 2019, 91, 8002\u22128007 8003 incompatible environment. The plasma setup used here is similar to the one previously described.12 Ionization waves were generated by applying a pulsed electric field E\u20d7 of high intensity (7.5 kV potential difference) in a He gas flow, within an external gas environment of 100% N2 flow in a surrounding tube (see Figure 1). The goal of such a setup is to better control the nature and yields of RONS in the plasma phase and subsequently in the exposed liquid phase, herein a physiological PBS (pH 7.4). In the following studies, only 100% N2 environment was used since this condition provides us with high concentrations of dissolved RONS at minute time-scale exposure durations. RONS derive from nitrogen excited species (e.g., NO\u00b0, N2 (FPS)) and oxygen species (e.g., HO\u00b0 and O) in the gaseous phase12 that were characterized by optical emission spectroscopy realized near the liquid surface (data not shown). As mentioned above, the RONS lifetime in aerated PBS at room temperature and physiological pH ranges typically from nanoseconds (HO\u00b0) to hours (NO2 \u2212), but major ones (NO\u00b0, O2\u00b0 \u2212, H2O2, ONOO \u2212) are stable in the second to minute time scale. Consequently, there is a need to detect and identify them in solution close to their source, i.e., the zone of interaction between the plasma and the solution (see scheme in Figure 1C). Electrochemical methods are particularly adequate to detect these reactive oxygen and nitrogen species.17 However, the basic principle of a three-electrode-based electrochemical cell is to apply a homogeneous and stable electric field (usually with a millivolt precision) between the WE and the REF. In the present experimental setup, a strong micropulsed electric field E\u20d7 of 10\u221220 kV\u00b7cm\u22121 18 amplitude and transient currents of about 1 A are propagated by the ionization wave and reach the solution where E\u20d7 strongly dissipates",
            " Herein, UMEs were fabricated from a platinum microwire insulated by a millimetric glass capillary (Figure S1). Then such minimal final size permits their positioning at short distances from the tip of the shielded container, and from the REF electrode (few millimeter distance). Therefore, this system limits the distance on which the potential difference is applied and obviously minimizes the impact of any polarizing-surrounding electric field.18,19 It even gives the possibility to insert two working electrodes next to the reference wire (Figure 1B) and achieve two different detections (see following paragraph). Eventually, as well-known in the field of electroanalysis, UMEs provide steady-state responses and fast response time, making these types of electrode ideal for monitoring concentration variations of redox species, which is the final goal of the setup development described herein. Besides, the two counter electrodes (CE) were placed in a separated shielded container (Figure 1B). Indeed, this separation is essential to avoid any interferences between redox species generated at both working and counter electrodes. Finally, the position of the electrode containers compared with the plasma spot was adapted (characteristic distance \u201cd\u201d in Figure 1A) and was maintained in the quartz cuvette using dedicated apertures. The final working setup is imaged in Figure 1C. In Situ Electrochemical Measurements during Plasma Exposure. The immunity of the electrochemical cell to the tension at the plasma generator and to electric field transported by ionization waves was assessed first by cyclic voltammetry (CV). Because of its working principle, CV should provide information on any overpotential or added electrochemical currents due to the plasma\u2212solution interaction. Voltammograms were then first acquired in a PBS solution, directly into the quartz cuvette where the plasma will be applied"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000056_j.precisioneng.2014.11.008-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000056_j.precisioneng.2014.11.008-Figure6-1.png",
        "caption": "Fig. 6. The meshing in FE analysis.",
        "texts": [
            " 5a, with an average constant axial stiffness. Frictional ontact between the spacer and spring was taken into account by nserting contact elements on the surface. Since the spring and spacer are mounted on the spindle at tight olerances, they can only move in axial direction. Therefore in he FE model, non-axial degrees of freedom of components were estrained using appropriate displacement boundary conditions. The finite element mesh was built using 3D brick elements, and he size of elements was chosen based on convergence study. Fig. 6 hows the finite element mesh used for this analysis. Once the FE odel was built, the external load on the system was applied in wo steps. First, an initial preload force was induced in the sysem by applying displacement boundary conditions on the back ace of the spring (see Fig. 5b), representing tightening of the claming nut. Then, the assembly was rotated at angular speeds ranging rom 1000 rpm up to 7000 rpm in steps to investigate the effects of entrifugal force on preload variation. The objective of the design was set to reduce the preload from ts high initial value of around 1500 N to a moderate value, around 00 N, at 7000 rpm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003775_fitee.1800570-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003775_fitee.1800570-Figure1-1.png",
        "caption": "Fig. 1 Schematic of a quadrotor unmanned aerial vehicle",
        "texts": [
            " The primary contributions of this paper can be summarized as: (1) A new AFTTCS is proposed for a QUAV against actuator faults with satisfactory performance and stability analysis is performed based on the Lyapunov stability theory; (2) In addition to actuator fault compensation, the proposed AFTTCS can achieve a satisfactory tracking performance when model uncertainties occur; (3) Actuator dynamics are considered to prevent actuators from exceeding their amplitudes and rate constraints, as well as to achieve a smooth actuator operation. As shown in Fig. 1, a QUAV is a multi-rotor aircraft with a simple geometric structure. Four motors are mounted on the midpoints of a rigid cross frame. The rotation of the fixed-pitch propellers produces thrusts for continuous QUAV flying. To balance the torques, two rotors (1 and 2) rotate in the clockwise direction, while the others (3 and 4) rotate in the opposite direction. The motions of a QUAV are controlled by changing the speed of each rotor. The pitch motion is achieved by conversely varying the speeds of the rotors (1 and 2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure3-1.png",
        "caption": "Fig. 3. Method to obtain the equivalent stiffness.",
        "texts": [
            " (1) The shaft segment is loaded, x k x i In this case the induced displacement \u03b4T on point P I ,k of the tooth I can be calculated as \u03b4T = x k x i \u00b7 r bS \u00b7 \u03c6T i (20) The influence coefficient f k,i for the displacement on point k due to the load acting on point i is calculated from the definition f T k,i = \u03b4T i / p i and F II, i = s p i , i.e., f T k,i = x k \u00b7 2 \u00b7 s \u00b7 r 2 bS \u03c0 \u00b7 (r 4 f \u2212 r 4 i ) \u00b7 G , (21) where s is the area of the discrete unit on point i . (2) The shaft segment is not loaded, i.e., x j > x i In this case, the twist displacement on each point P j in the segment is equal to the deformation caused by the reaction force acting on point P i . Similarly, the influence coefficient f T j,i is f T j,i = x i \u00b7 2 \u00b7 s \u00b7 r 2 bS \u03c0 \u00b7 (r 4 f \u2212 r 4 i ) \u00b7 G . (22) Because the planet shafts are assembled with the carrier ( Fig. 3 ), the deformations of the two types of components are coupled with each other. A load acting on a tooth of a planet gear can not only cause deflection of the planet shaft (called \u201cdeflection type 1\u201d), on which the planet is mounted, but also the deformation of the carrier, which causes further the deflection of the other planet shafts (called \u201cdeflection type 2\u201d). It is thus convenient to derive an influence coefficient from the combined deformations of the planet shaft and the carrier. The deflection of the planet shaft, which is determined according to the integration equation, \u03b4(z) = \u222b \u222b M(z) E 1 I(z) d zd z + c 1 \u00b7 z + c 2 , (23) can be solved by using the method of singularity function (see Appendix ) under different boundary conditions considering the deformation influences of the carrier"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000072_aim.2014.6878331-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000072_aim.2014.6878331-Figure9-1.png",
        "caption": "Fig. 9. Optimum workpiece placement",
        "texts": [
            " The white areas can not be reached by the robot because of its joint limits. It is noteworthy that 0xOW , 0yOW , 0zOW , Q4 and \u00b5 are the only decision variables considered by the ga and fmincon functions in this optimization problem solving. As a matter of fact, an optimal \u03b2 vector is searched at each iteration of the genetic algorithm and at each iteration of the interior-point algorithm. This vector is obtained in such a way that it minimizes the objective function fMQC and avoids discontinuities in the robot joint space along each segment. Figure 9 represents the optimum workpiece placement found by solving optimization problem (26). It is noteworthy that the worst workpiece placement is obtained by maximizing the objective function fMQC while respecting the constraints of optimization problem (26). Figure 10 illustrates the tool displacement ci with respect to the tool path point number and \u03b2 angle for the worst workpiece placement. The green curve characterizes the optimum redundancy planning scheme, i.e., \u03b2opt vector, whereas the red curve represents the worst redundancy planning scheme for this workpiece placement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002351_s12206-015-0506-2-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002351_s12206-015-0506-2-Figure7-1.png",
        "caption": "Fig. 7. Appearant standing wave at 280 km/h.",
        "texts": [
            " To further validate the proposed tire model, a high(speed simulation of the tire is performed. When a tire rolls with speeds higher than its critical speed, intensive waves are gen( erated that starts from the tire\u2019s leading edge and travels around the circumference of the tire. At a certain speed, to an observer, these waves appear not to be moving, therefore it is known as the standing wave phenomenon [19, 20]. The high( speed simulation of the proposed tire model demonstrated that the standing wave for the truck tire starts to develop approxi( mately at 280 km/h as shown in Fig. 7. These results are in close agreement with the actual test measurements. A methodology to obtain a condensed finite element based tire model was presented. In the process of formulating the proposed tire model, several issues in model reduction of the tire were addressed. The presented methodology is based on the rotation(invariant characteristic of the tire which leads to introducing a non(rolling mesh. A mixed Eulerian(Lagrangian description was used to derive the equations of motion. A well( known component mode synthesis method was used to con( struct a condensed tire model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001649_j.sna.2018.10.031-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001649_j.sna.2018.10.031-Figure1-1.png",
        "caption": "Fig. 1. The diagram of the CCD-based tracking system.",
        "texts": [
            " The rest of this paper is organized as follows: Section 2 presents a rief explanation of the CCD-based tracking system, the traditional ual closed-loop method for the CCD and MEMS accelerometer ased the low-cost tracking system, and the implementation of the asic acceleration feed-forward. Section 3 focuses on the acceleraion feed-forward based on data fusion of model output and sensor ata. Section 4 implementation of acceleration feed-forward conroller is discussed. Section 5 the experiments are set up to verify he effectiveness of the proposed methods. Concluding remarks are resented in Section 6. . Implementation of the basic acceleration feed-forward .1. CCD-based tracking system The diagram of the CCD-based tracking system is shown in Fig. 1. hen the system is tracking a target, the phenomenon that the taret remains at the center of the target surface is regarded as ideal racking. The angle between the target and the center of the taret surface is called the LOS error and is detected by the CCD. The OS error is provided to the system controller to achieve the posiion closed-loop tracking. The tracking system in Fig. 1 is a two-axis imbaled system. A stands for the azimuth axis and E represents the levation axis. Description of E axis is emphasized due to the simiarity of two axes. E target is the LOS error, E platform is the platform ngular position, E target is the target trajectory. The relationship mong them is given by Eq. (1): E target = E target + E platform. (1) To achieve the feed-forward control, the target trajectory is equired. The system can only get the LOS error from the CCD, so extra sensors are necessary to provide the platform\u2019s angular position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003973_icieam.2019.8743076-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003973_icieam.2019.8743076-Figure3-1.png",
        "caption": "Fig. 3. Temperature distribution inside the induction motor, at t = 8000 s, considering a healthy motor condition.",
        "texts": [
            " Due to the large data sets produced while running the thermal simulations, the analysis of the simulation results will solely focus on the analysis of the temperature distribution inside the motor at the instants in which the thermal steady state is reached. According to IEEE and IEC standards, the thermal steady state of electrical machines is reached when the temperature increment along one hour is less than 2 \u00b0C [11], [12]. In this context, an analysis based on the observation and relative comparison of scenarios will be employed. Fig. 3 to Fig. 6 depict the temperature distribution inside the motor, in \u00b0C, considering the motor operation under healthy and faulty conditions. Note that the temperature distributions presented in Fig. 3 to Fig. 6 refer to the instant of time at which the thermal steady state condition was already met. In the conditions of load torque and initial room temperature defined for the simulations, the thermal steady state conditions are observed before the instant t = 8000 seconds. As depicted in Fig. 3, there is a consistent and uniform distribution of temperatures inside the motor, typical of a healthy operation scenario. As expected, the rotor is the warmer part of the motor, followed by the stator. The rotor temperature is indeed quite uniform. Inside the stator, the windings are slightly warmer than the rest of the stator elements. Meanwhile, it is also observed that the temperature of the stator iron smoothly decreases in the direction of the stator outer boundaries. Fig. 4 depicts the temperature distribution map inside the motor for the scenario in which a short-circuit fault between 6 turns of phase W is imposed. The short-circuit resistance (1.35 ) is adjusted so that the short-circuit current flowing through the short-circuit branch equals the motor rated current. When comparing the results of Fig. 4 with the healthy condition (Fig. 3), it is observed that there is a general increment of the motor temperature. All motor components experience an increment of temperature of up to 10 \u00b0C. The rotor remains as the warmer component of the motor. As a consequence of the short-circuit fault, the temperature distribution along the stator is not uniform any longer. The slots that host the faulty turns are 2 - 4 \u00b0C warmer than all the other stator slots. The neighboring slots also suffer a minor temperature increment. The progressive decrement of temperature in the direction of the stator outer boundary is maintained in this fault condition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001568_j.mechmachtheory.2018.07.009-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001568_j.mechmachtheory.2018.07.009-Figure4-1.png",
        "caption": "Fig. 4. Schematic representation of the bi-planar Stewart platform manipulator. (a) General bi-planar Stewart platform manipulator; (b) Leg architecture.",
        "texts": [
            " The SPM chosen in this work is one with two irregular hexagons serving as its fixed base and moving platforms, respectively. The dimensions of these platforms and other physical parameters, such as the mass and inertia properties of the manipulator, are listed in Appendix B . The derivation of the kinematic constraints as well as the equation of motion are described in the following sections. The geometry of a general bi-planar Stewart platform manipulator along with its detailed leg architecture is schematically represented in Fig. 4 . The platforms are connected by six U P S (universal-prismatic-spherical) legs, of which the prismatic joint is actuated in each leg. The angular displacements at the i th universal joint are denoted by \u03c8 and \u03c6 . The length of the i i fixed and the moving part of each leg is denoted by l b i and l a i , respectively, the total leg length at any point of time being l i . The centres of mass of the fixed and moving parts of the legs are located at p b i and p a i , respectively. The location of the spherical joint of each leg connecting to the moving platform is denoted by p i , in the global frame of reference o B - X B Y B Z B ",
            " One such configuration, where configuration-space singularity occurs for a five-bar manipulator of architecture: l = r = 1 , l = l = r = 2 is \u03b8 = 0 and \u03b8 = 0 , as shown in Fig. 2 . 1 1 0 2 2 1 2 The values of the architecture parameters for the SPM have been adopted from Shanker and Bandyopadhyay [24] . The positions t = [ t x , t y , t z ] and b = [ b x , b y , b z ] , respectively, of the vertices of the moving and fixed platforms measured with respect to their local frame of reference (see, e.g., Fig. 4 ) are listed in Table B.8 . The components of the manipulator are assumed to be made of steel, with a density \u03c1 = 7900 kg/m 3 . The legs are assumed to be of cylindrical shape (see Fig. 4 b). The fixed part of each leg is assumed to be a hollow cylinder, while the moving part is a solid cylinder. The masses of the fixed parts of the legs are given by: m b i = \u03c1\u03c0 l b i (D 2 \u2212 d 2 ) / 4 , where D and d represent the outer and inner diameters, respectively. The inertia matrices of the fixed parts are computed from the standard formula for hollow cylinders and are given by: I b i = diag ( m b i 48 ( 3(D 2 + d 2 ) + 4 l 2 b i ) , m b i 48 ( 3(D 2 + d 2 ) + 4 l 2 b i ) , m 8 (D 2 + d 2 ) ) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure9-1.png",
        "caption": "Fig. 9. The section view of a three-stage planetary reducer for analysis.",
        "texts": [
            " The distributed contact stress on flanks and the load sharing among planets, obtained from the proposed LTCA model, are at first compared with those from the FEM software (MSC.Marc). The effects of the deformations of the sun gear, the carrier with the planet shafts on the distribution of the contact stress are then identified using the LTCA approach. The variation of shared loads and contact stresses of an individual single tooth pairs, as well as the contact pattern at specific contact positions are analyzed and discussed later. The final stage of a three-stage planetary reducer (see Fig. 9 ) is chosen as the study case for analysis. The parameters of the planetary gear set are listed in Table 4 . Five planet gears are separated equally around the sun gear to carry out the high torque. The tooth profile is assumed as exact involute without any flank modification in the case. The equivalent stiffness of the carrier is obtained from FEM result by using CAD software Autodesk Inventor, see Fig. 10 . The total load with a value of 453.26 kN is applied either on the input side or on the output side of the carrier"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002720_978-3-319-24055-8-Figure5.3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002720_978-3-319-24055-8-Figure5.3-1.png",
        "caption": "Fig. 5.3 Studied gearbox for the dynamic response computation",
        "texts": [
            " Indeed, dispersion of manufacturing errors generates a strong variability of the dynamic behavior and noise radiated from geared systems (sometimes up to 10 dB [9, 10]). 5 Vibro-Acoustic Analysis of Geared Systems \u2026 65 A statistical study of solutions permits to have a good overview of how the solution can be deteriorated when the manufacturing errors (dispersion over the optimization parameter values) and assembling errors (lead summed up and involute alignment deviations, respectively fHb and fga) are considered. The dynamic response computation procedure is applied to an automotive gearbox displayed in Fig. 5.3. This computational scheme requires a finite element model of the complete gearbox in order to obtain its modal basis. The contact between the gears is modelled with a stiffness matrix linking the degrees of freedom of each pair of meshing gears. To achieve that, the mean value of the mesh stiffness is taken, leading to mean modal characteristics. The parametric mesh stiffness k\u00f0t\u00de isdirectly related to the applied torque T and the static transmission error STE\u00f0t\u00de with: k t\u00f0 \u00de \u00bc 1 Rb @T @STE\u00f0t\u00de \u00f05:2\u00de The scheme uses then a powerful resolution algorithm in frequency domain to solve the dynamic equations with an iterative procedure [11, 12]",
            " The comparison is based on predominance of orders and modes, in terms of frequency and amplitude. The dominant orders and the frequency ranges exhibiting a dynamic amplification correctly determined. An order tracking has also been done in order to compare properly the vibration measurements with the computations. 74 A. Carbonelli et al. The first and second orders of the two meshes have been considered (the first mesh corresponds to Z1/Z2 = 35/39. The second mesh corresponds to Z3/Z4 = 16/69 as specified in Fig. 5.3). The acceleration of one housing point for the second order of the second mesh is displayed in Fig. 5.13. The dynamic model has been tuned in different operating conditions explaining some non-negligible frequency shifts and modal response differences. However the agreement between the measurements and the computations remains satisfying for a predicting tool. On the contrary to the measurements, the simulation can take into account the variability of the results. Extracted from teeth metrology, a dispersion study has been performed to determine the envelope of the dynamic response"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001846_978-90-481-9707-1_116-Figure11.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001846_978-90-481-9707-1_116-Figure11.1-1.png",
        "caption": "Fig. 11.1 Cross sections of five common beams",
        "texts": [
            " The aim of carrying out the single quadrotor arm analysis is to provide a useful summary on the performance of different quadrotor arm shapes, and then among the few possible candidates under the constrained of weight and size, a best solution in terms of shape, length, and dimension can be obtained. In this analysis, the quadrotor arm is approximated as a cantilever beam which has a fixed end and a free end. Five different types of cantilever beams are designed and analyzed. The cross sections for each of the beams are shown in Fig.11.1. A few different parameters on the dimension of the beams will be varied, and the corresponding natural mode (1\u20134) will be compared. Note that the composite material for all beams to be simulated has the properties of carbon fiber code name carbon/epoxy T300/976. All simulations are done in Nastran by employing the discrete element with six degrees of freedom (DOF) per node, and all nodes at one of the tips are assumed to be fixed by setting all six DOF to zero. The first variable to be investigated is the length of the beam. It is well known that slender bodies are more easily exposed to vibration, or in other words, shorter beams are stiffer. To verify the relationship between the length of the beam and its natural modes, the rectangular beam (first cross section in Fig. 11.1) with different length is analyzed in the simulation. In Table 11.1, a summary of the Nastran natural mode analysis results obtained for this study is given. Based on the results, it is evident that natural frequencies for the first four modes increase as the structure becomes shorter. In general, beams of other shape show similar behavior, and thus the results are trivial and not to be included here. Despite the results favoring shorter beams, there are other restrictions on the minimum length of the quadrotor arms",
            " One important factor would be the aerodynamic interferences between the rotors, which generally limits the minimum length of the quadrotor arms to be at least twice the rotor radius. Length Natural frequency (Hz) (mm) Mode 1 Mode 2 Mode 3 Mode 4 The second variable to be investigated is the thickness of the material. As two common thickness of carbon fiber sheet or beam available commercially are of 0.5 and 1 mm thick, the natural mode of all five general shapes of beams are analyzed and compared in 0.5 and 1 mm thickness. For fair comparison, the model for each cross section was constructed with length l D 60mm, width w D 6mm, and height h D 6mm (see Fig. 11.1 for illustration). Table 11.2 shows the comparison between beam with thickness t D 1mm and t D 0:5mm, together with their calculated weight. One can notice that in general, the beam having a closed cross-sectional shape, i.e., the rectangular hollow and circular hollow beams, has a much higher natural frequency compared to other shapes. Also, the thickness of the material has little effect on them, while the weight could be half as light. The last variable to be investigated is the width and height of these closed crosssectional beams"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000369_iros.2013.6696872-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000369_iros.2013.6696872-Figure1-1.png",
        "caption": "Fig. 1. The simple bipedal model consists of two rigid legs of length \u2113 and mass m. The leg mass has an arbitrary distribution with no restriction on the location of its center of mass. The prismatic actuator on the stance leg (not shown in the figure) can apply the extensional force F along the leg and the revolute actuator at the hip can apply torque \u03c4 between the legs.",
        "texts": [
            " Then, the energetics of each impulse is calculated using a work-based energetic cost model and through the so-called \u2018overlap parameter\u2019 [8] that quantifies the order and percentage overlap of the retraction torque and the push-off force. Finally, given the push-off and retraction impulses, the optimal relative timing of push-off and swing retraction is calculated by minimizing the net energetic cost of walking. The result is valid for both periodic and aperiodic gaits. The first model used in this paper is shown in Fig. 1. It consists of two identical rigid legs, each with length \u2113 and mass m. The leg mass is arbitrarily distributed with no restrictions on its inertia or the location of its center of mass. This biped is powered with two actuators: (i) a prismatic actuator on the stance leg (not shown in Fig. 1), applying the extensional force F along the leg, and (ii) a revolute actuator at the hip, applying the torque \u03c4 between the legs. In general, these actuators can have many arbitrary force/torque profiles (as long as walking is feasible), where each combination results in a different energetic cost. However, the energy optimality requirement restricts our choice. To identify the optimal profiles for leg force and hip torque in our model, consider the following observations from both human experimental data and previously done numerical optimizations",
            " The impulsive push-off force and the impulsive retraction torque are quantified by their net impulse denoted by P and R. If we denote the time instant just before both impulsive actions by t\u2212pr and the time instant just before heel-strike (just after both impulsive actions are completed) by t\u2212h , the push-off impulse P and the retraction impulse R are given by P = \u222b t\u2212 h t\u2212pr F (t) dt (1) R = \u2212 \u222b t\u2212 h t\u2212pr \u03c4(t) dt. (2) We assume that the sign of the push-off force and the retraction torque does not change during their application. Therefore, based on our convention in Fig. 1 for the positive directions, the push-off force F is always positive, and the retraction torque \u03c4 is always negative. Hence, P > 0 and R> 0. Although the impulsive push-off and retraction actions take place at the same biped configuration (Fig. 2a), we treat P and R as isolated in time (one after the other) or as having some (specified) overlap with each other (to be discussed later in detail). We will show in section IV that the relative timing of these two impulses can change the energetics of the gait",
            " at t\u2212h . This is possible if the retraction impulse R is large enough to push the leg backwards so that the leading foot is moving downwards at t\u2212 h . To find the joint velocities at t\u2212h (just after both impulsive actions) in terms of the kinematic variables at t\u2212pr and the impulse magnitudes P and R, we integrate the biped\u2019s equations of motion (EoM) over the infinitesimal interval (t\u2212pr, t \u2212 h ). Throughout the impulsive push-off and retraction the model dynamics follow the EoM in single stance (Fig. 1) which can be obtained from (i) linear momentum balance equation of the whole mechanism along the stance leg, (ii) angular momentum balance (AMB) equation of the swing leg about the hip joint, and (iii) AMB equation of the whole mechanism about the stance foot. After rearrangement, these three equations can be written in the following standard form: M(q) q\u0308+ h(q, q\u0307) =   F \u03c4 0   , (3) where the vector q(t)= [ \u2113(t), \u03c6(t), \u03b8(t) ]T determines the biped configuration, M is the symmetric mass matrix, the vector h includes the Coriolis, centrifugal and gravity terms, and the hip torque \u03c4 and the stance leg force F have impulsive profiles in t\u2212pr6 t6 t\u2212h with impulse magnitudes given by (1) and (2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001157_j.precisioneng.2017.07.013-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001157_j.precisioneng.2017.07.013-Figure2-1.png",
        "caption": "Fig. 2. Schematic of the new assist gas injection system. (Positioning parameters: X \u2013 d b",
        "texts": [
            " The main drawback of this approach s the low mass flow entering into the kerf [41]. This reduction in ass flow decreases the removal of molten material, and the cut uality is compromised. It seems reasonable that the most suitable shape of the assist gas et should be rectangular to avoid the precedent drawbacks. In this ay, the jet can naturally adapt to the geometry of the cut kerf, and he choking would be even smaller. Furthermore, the jet can be thin n one side (to mimic the kerf width), but large on the other side o avoid the reduction in mass flow (see Fig. 2). In this regard, Son t al. studied the impingement characteristics of jets from off-axial upersonic rectangular nozzles with a simulated cut kerf using the chlieren technique, and Pitot pressure measurements [42]. They ompare the results with those previously observed for convergng coaxial nozzles. They found a marked reduction in the strength f the MSD, and higher Pitot pressures on the kerf surface. These esults suggest a favourable aerodynamic interaction of this kind f jets during laser cutting. According to these considerations, we developed a cutting head ith a supersonic nozzle with a rectangular cross-section (i.e. on-axisymmetric) to inject the assist gas into the cut kerf. Fig. 2 epicts a schematic drawing of this experimental setup. The superonic rectangular nozzle was designed in accordance with the basic rinciples of the steady one-dimensional compressible flow with riction. There are three main positioning parameters for the jet, as epicted in Fig. 2. These are: the distance between the cut front and he impinging point of the gas jet (X), the stand-off distance from he supersonic nozzle to the workpiece (Z), and the angle between he laser beam axis and the axis of the gas jet ( ). They were adjusted y means of cutting experiments to obtain the higher optimum emoval of molten material by the gas jet. The results have been ompared with those obtained using a converging nozzle (coaxal with the laser beam and with axisymmetric geometry), and an jet exit static pressure, P0: ambient static pressure, P/P0: pressure ratio, Me: Mach number at the nozzle exit, Ae/At: exit-to-throat area ratio, and De: width of the jet in the nozzle exit)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002898_asjc.1411-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002898_asjc.1411-Figure1-1.png",
        "caption": "Fig. 1. The PVTOL aircraft.",
        "texts": [
            " = 2l cos \ud835\udefc J F , (1) where (x, y) is the position of the center of mass of the aircraft, \ud835\udf03 is the roll angle of aircraft with respect to the horizon, T and F are two control inputs, respectively standing for the thrust directed out the bottom of the aircraft and the rolling moment produced by a couple of equal forces, whose direction is not perpendicular to the horizontal body axis, but tilted by some fixed angle \ud835\udefc \u2208 (\u2212\ud835\udf0b\u22152, \ud835\udf0b\u22152), M represents the mass of the aircraft, J the moment of inertia about the center of mass, 2l the distance between the wing-tips, and g the gravitational acceleration (Fig. 1). Assumption 1. The modeling parameters (M, J, l, \ud835\udefc) and the gravitational acceleration g are unknown constants with each one owning known upper/lower bounds, denoted by subscripts M\u2215m, that is 0 < Mm \u2264 M \u2264 MM , 0 < Jm \u2264 J \u2264 JM , 0 < lm \u2264 l \u2264 lM , |\ud835\udefc| \u2264 \ud835\udefcM < \ud835\udf0b\u22152, 0 < gm \u2264 g \u2264 gM . Define the position error as x\u0304 = x \u2212 xd , y\u0304 = y \u2212 yd , (2) where (xd , yd) is the desired constant position. The stabilization control problem of PVTOL aircraft studied in this work is stated as: under Assumption 1, find a control law (T(\u22c5),F(\u22c5)) to steer the position error, the roll angle and the corresponding velocities to zero, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002267_tmag.2013.2281460-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002267_tmag.2013.2281460-Figure6-1.png",
        "caption": "Fig. 6. 2-D computational model of motor. (a) Whole model. (b) Meshing.",
        "texts": [
            " From this figure, we can see that B locus at point M located at the T-joint part of transformer core shows elliptical shape clearly when compared with that at points L and R no matter using the proposed model or E&S model, and the simulated loci are all in good agreement with the measured ones. To illustrate the effectiveness of the proposed model further, a 5.5 kW three-phase induction motor model is taken as another example, in which the pole number is two, the stator slot number is 30 and rotor is 26, as shown in Fig. 6(a). The motor runs at no-load and to make the mesh of stator fixed when rotor rotates an edge line at air region is defined, as shown in Fig. 6(b). The B loci distributions in stator core calculated from the proposed complex-valued model are shown in Fig. 7. From this figure, it is can be observed that the NGO electrical steel is magnetized with a lot of rotational magnetic flux densities, which are different from the transformer core. Fig. 8 shows the iron loss distribution in stator core. In Fig. 8, the distribution rules calculated from proposed complex-valued and E&S models in [8] are almost same, in which the maximum loss appears at the bottom of the slot near the middle part of stator core where the magnetic force line concentrates and goes along the transverse direction",
            " The total stator core losses calculated using proposed complex-valued and E&S models in [8] are 92.3 and 94.6 W, respectively, which are different slightly. According to (3) and (10), we know that the loss value has relationship with the hysteresis loop of electrical steel at both rolling and transverse directions. To explain the difference, we calculate the hysteresis loop at rolling and transverse directions in each mesh element and Fig. 9 shows the hysteresis loop of rolling direction at points a\u2013d in Fig. 6(a). From Fig. 9, we can observed that the hysteresis loops using two methods mentioned above are not completely overlapped because the complex-valued model employs the assumption of magnetic field quantities changing with sinusoidal waveforms. To check the simulation accuracy, a prototype of induction motor was made, as shown in Fig. 10, in which eight test coils were winding at the bottom of stator slots. The measurement data of amplitude of magnetic flux are compared in Fig. 11 with simulated ones, which include the results from the orthogonal B\u2013H curve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001084_j.mechmachtheory.2017.05.017-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001084_j.mechmachtheory.2017.05.017-Figure2-1.png",
        "caption": "Fig. 2. Coordinate systems applied for tooth contact analysis.",
        "texts": [
            " \u21c0 r t1 and \u21c0 n t1 are the position and unit normal vectors of a rack-cutter, respectively, both determined by Eq. (2) . \u23a7 \u23aa \u23a8 \u23aa \u23a9 \u21c0 r t1 ( u 1 , l 1 ) = [ M] t1 ,a 1 [ u 1 , 0 , l 1 , 1] T \u21c0 n t1 = \u2202 \u21c0 r 1 \u2202 u 1 \u00d7 \u2202 \u21c0 r 1 \u2202 l 1 / \u2223\u2223\u2223\u2223 \u2202 \u21c0 r 1 \u2202 u 1 \u00d7 \u2202 \u21c0 r 1 \u2202 l 1 \u2223\u2223\u2223\u2223 (2) where, u 1 and l 1 are the surface parameters of the rack cutter; [ M ] t 1, a 1 is the 4 \u00d7 4 matrix that describes the coordinate transformation from S a 1 to S t 1 . A pair of regular surfaces being in contact continuously means that they are always in tangency over time without separation and interpenetration. As shown in Fig. 2 , to perform TCA, the input pinion tooth face 1 rotates with an independent variable \u03d51 , and the output gear tooth face 2 is driven to rotate with a dependent variable \u03d52 . The coordinate system S f is connected rigidly to the gear housing. When surfaces 1 and 2 are in tangent, their position vectors \u21c0 r (i ) f (i = 1 , 2) and unit normal vectors \u21c0 n (i ) f (i = 1 , 2) considered in S f must be equal, which can be represented by Eqs. (3 ) and (4) . { \u21c0 r ( 1 ) f ( u 1 , l 1 , L 1 , \u03d5 1 ) = \u21c0 r ( 2 ) f ( u 2 , l 2 , \u03d5 2 ( \u03d5 1 )) \u21c0 n ( 1 ) f ( u 1 , l 1 , L 1 , \u03d5 1 ) = \u21c0 n ( 2 ) f ( u 2 , l 2 , \u03d5 2 ( \u03d5 1 )) (3) \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 \u21c0 r ( 1 ) f ( u 1 , l 1 , L 1 , \u03d5 1 ) = [ M] f 1 \u21c0 r 1 ( u 1 , l 1 ) \u21c0 n ( 1 ) f ( u 1 , l 1 , L 1 , \u03d5 1 ) = [ L ] f 1 \u21c0 n 1 ( u 1 , l 1 ) \u21c0 r ( 2 ) f ( u 2 , l 2 , \u03d5 2 ) = [ M] f 2 \u21c0 r 2 ( u 2 , l 2 ) \u21c0 n ( 2 ) f ( u 2 , l 2 , \u03d5 2 ) = [ L ] f 2 \u21c0 n 2 ( u 2 , l 2 ) (4) where [ M] f i ( i = 1 , 2 ) is the 4 \u00d7 4 matrix that describes the coordinate transformation from S i to S f , and [ L ] f i ( i = 1 , 2 ) is the 3 \u00d7 3 submatrix of [ M] f i ( i = 1 , 2 ) ; u i and l i are the parameters of the pinion ( i = 1) and gear ( i = 2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure5-1.png",
        "caption": "Fig. 5. Boundary fillet \u2014 type IIb of the rack-cutter.",
        "texts": [
            " (4) and taking into consideration that _X\u03b7 \u00bc 2r sin2\u03c6; \u20acX\u03b7 \u00bc 4r sin\u03c6 cos\u03c6; _Y \u03b7 \u00bc \u22122r sin\u03c6 cos\u03c6; \u20acY \u03b7 \u00bc \u22122r cos2\u03c6; \u00f07\u00de equation of the radius of the curvature of the curve \u03b7, the following formula is obtained \u03c1\u03b7 \u00bc 4r2 sin4\u03c6\u00fe 4r2 sin2\u03c6 cos2\u03c6 3=2 8r2 sin2\u03c6 cos2\u03c6\u22124r2 sin2\u03c6 cos2\u03c6 \u00de \u00f08\u00de in, after a transformation, appears as follows: \u03c1\u03b7 \u00bc 2r sin\u03c6 \u00bc mz sin\u03c6: \u00f09\u00de 3.2. Boundary fillet \u2014 type IIb (curve \u03be) It is obtained as a trajectory of the point b of the coordinate system XbObYb connected with the gear (Fig. 5), drawn in a coordinate system X\u03beO\u03beY\u03be, connected with the rack-cutter. The obtained trajectory \u03be, as it was already explained, represents a shortened cycloid, whose parametric equations can be written as follows where and th X\u03be \u00bc r\u03c6\u2212rb sin\u03c6 \u00bc X\u03be \u03c6\u00f0 \u00de ; Y\u03be \u00bc \u2212r \u00fe rb cos\u03c6 \u00bc Y\u03be \u03c6\u00f0 \u00de ; \u00f010\u00de rb is the radius of the base circle, defined by the formula rb \u00bc r cos\u03b1 \u00bc 0:5mz cos\u03b1 : \u00f011\u00de In the initial position of the coordinate systems, when at \u03c6=0, point b coincides with point b\", and at \u03c6=\u03b1 \u2013with point b'\u2261A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003930_jmems.2019.2903457-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003930_jmems.2019.2903457-Figure2-1.png",
        "caption": "Fig. 2. (a) Fabrication scheme of sensor electrode on polyimide substrate; (b) optical images of sensor electrodes after step (ii) as seen from the backside. The Cr region can be seen through the translucent polyimide; (c) Sensor electrode top view after step (iii).",
        "texts": [
            " Without such a layer it was experimentally observed that the working and counter electrodes wrinkle and then physically deform after a very short duration of functionality; further, the absence of an adhesion layer in the pad region rendered subsequent processing more difficult (e.g., delamination of electroplated copper in the contact pads during soldering). To overcome these difficulties, we developed an electrode fabrication scheme that incorporates a partial adhesion layer and a mesh electrode morphology as shown in Fig. 2. An adhesion layer of chromium is patterned only in the contact pad region, where copper will be electroplated. The mesh structure consists of 20-\u00b5m-diameter holes arranged in a radial array. The first thin layer of Au/Cr was deposited onto the polyimide substrate by electron-beam evaporation and lithographically patterned by lift-off, followed by the deposition and patterning of a second, thicker Au layer. To achieve the mesh design, the working and counter electrodes were patterned with an excimer laser, whose micromachining conditions were optimized to ablate the thin film metal without damaging the underlying polymeric substrate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003416_15325008.2018.1444689-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003416_15325008.2018.1444689-Figure2-1.png",
        "caption": "FIGURE 2. Five-phase standalone wind generator based on the information available from GWEC and MNR 2015-2016.",
        "texts": [
            " The effect of modulation index is studied and the optimum modulation index is identified for different wind speeds. The reliability of the machine under fault condition is examined with one or two phases open. The results are compared with the three-phase generator Wind turbine drives the FPIG through gearbox. The five-phase AC output of the generator is rectified and filtered and feeds a five-phase inverter. It converts the variable dc output into five-phase ac as required by the load. A five-phase capacitor bank is used for self-excitation of the induction generator Figure 2 represents the general block diagram of standalone five-phase wind generator. The mathematical modeling of the five-phase wind generator components namely wind turbine, FPIG, five-phase rectifier and five-phase inverter are discussed in the following sections. The output power generated by the wind turbine is given by Ptur = 0.5\u03c1ACp(\u03bb)Vw 3 Watts (1) where \u03c1 = air density (kg/m3) A = area swept out by the turbine blades (m2) R = radius of the wind turbine rotor (m) Vw = wind velocity (m/s) Cp, a dimensionless power coefficient depends on the wind velocity and constructional characteristics of the turbine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003482_s00170-018-2324-z-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003482_s00170-018-2324-z-Figure3-1.png",
        "caption": "Fig. 3 Finite element model of oblique cutting",
        "texts": [
            " The cutting environment is dry cutting. In the cutting tests, the tools were ground into different geometrical angles. The detailed cutting parameters are listed in Table 3. After the oblique cutting of TC21 alloy, the residual stress of cutting layer was measured by the X-ray residual stress tester. The test location of residual stress is shown in Fig. 2. In Fig. 2, four points on the cutting surface of the sample after the oblique cutting were measured. The 3D oblique finite element model of cutting for TC21 alloy is shown in Fig. 3. The model was established by the software ABAQUS 6.10. The workpiece\u2019s dimension is 2 mm \u00d7 2 mm \u00d7 10 mm. The angle of inclination is 20\u00b0. The mesh of machined layer is refined and the mesh type is C3D8RT (An 8-node thermally coupled brick, trilinear displacement, temperature, reduced integration, and hourglass control.), and the tool mesh type is set to be C3D4T (a 4-node thermally coupled tetrahedron, linear displacement, and temperature). The cutting environment is dry cutting and the initial environment temperature is set to be 20\u00b0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000135_icra.2012.6225025-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000135_icra.2012.6225025-Figure2-1.png",
        "caption": "Fig. 2. Section frames",
        "texts": [
            " Two adjacent segi\u22121 and segi are connected tangentially at the connection point pi\u22121 as shown in Fig. 978-1-4673-1405-3/12/$31.00 \u00a92012 IEEE 4340 2.(a), i.e., the two sections share the same tangent at pi\u22121. For clarity, we consistently use black, red, and green colors to draw section 1, section 2, and section 3 of the OctArm in this paper. Each section i, i = 1, 2, 3, has three degrees of freedom that can be directly changed by the OctArm actuators [12], which are controllable variables: curvature \u03bai, length si, and orientation angle \u03c6i from yi\u22121 axis to yi axis about zi axis. Fig. 2.(b) shows one example segi, its frame, and controllable variables. Note that the center of ciri, ci, always lies on the xi axis, with \u22131/\u03bai being the x coordinate in the i-th frame, where \u03bai is the curvature. Note also that ci lies on the positive xi axis if \u03bai < 0 and on the negative xi axis if \u03bai > 0. When \u03bai = 0, segi is a straight-line segment along the z axis, and ci can be considered at either + or \u2212 infinity along the x axis. The configuration of the entire arm is determined by the control variables of each section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001377_j.compeleceng.2018.02.012-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001377_j.compeleceng.2018.02.012-Figure5-1.png",
        "caption": "Fig. 5. Correspondence between the DCS, the SCS and the CCS.",
        "texts": [
            " The coordinates ( x, y, z ) of PC are deduced from the longest sub list L sp [ a ] taken from among the L s [ a ] extracted lists. A 2 D representation of L sp [ a ]\u2019s clusters is similar to the one of Fig. 4 ; they are all from corona a . Here, we must redirect the flow by breaking this long path at 2/3 of its length. For that purpose, we must place an actuator below the cluster situated at 2/3 of the chain; from there, to optimally reduce the height of tree T , the flow of data should circulate as shown in Fig. 5 . Let ( a, v 0 , h 0 ) be the coordinates of the cluster situated at the 2/3 point of list L sp [ a ]. The coordinates ( x, y, z ) of PC are x = a \u2013 1; y = v ; and z = h . As long as there are available actuators, the sink should move an actuator in clusters ( x, y, z ), 0 0 Please cite this article as: J.F. Myoupo et al., Fault-tolerant and energy-efficient routing protocols for a virtual three-dimensional wireless sensor network, Computers and Electrical Engineering (2018), https://doi.org/10",
            " The Z-axis is chosen such that it coincides with the right border of the first horizontal angular section (horizontal angular Section 0). The X-axis is chosen such that it coincides with the right border of the first vertical angular section (vertical angular Section 0) and this axis is at a quarter turn (trigonometric direction) from the Z-axis . The Y-axis is at a quarter turn from the Y- and Z-axes so that the reference ( sink, \u2212\u2212\u2212\u2192 sinkX , \u2212\u2212\u2212\u2192 sinkY , \u2212\u2212\u2212\u2192 sinkZ ) is direct (as in Fig. 3 a). When we extract this reference, we have the SCS of Fig. 5 a. Any point M of the DCS is discoverable in the SCS using p (its distance from the sink), \u03b8 (angle measured from the X-axis , 0 \u2264 \u03b8 < 2 \u03c0 ), and either \u03d5 (angle measured from the Z-axis , 0 \u2264\u03d5 < \u03c0 ) or \u03b4 (angle measured from the X- axis, \u2013 \u03c0 /2 \u2264 \u03b4 < \u03c0 /2 ) as in Fig. 5 a. Nevertheless, \u03b4 and \u03b8 are linked by the relationship \u03b4 = \u03c0 2 \u2212 \u03d5. Similarly, any point M of the DCS or SCS is also discoverable in the CCS using its coordinates ( M x , M y , M z ), as shown in Fig. 5 b. By letting e denote the thickness of a corona, \u03b1 the angle of a vertical angular section and \u03b2 the angle of a horizontal angular section, we can state Corollaries 1 and 2 . Corollary 1. Let M be a sensor of the cluster (c, v, h) assumed at its center. Let \u03b8 \u2032 = v \u00d7 \u03b1 + \u03b1 2 and \u2019 = h \u00d7 \u03b2 + \u03b2 2 . In the SCS, M has the coordinates (p, \u03b8 , \u03d5) given by the following: i f (\u03d5 \u2032 < \u03c0) { p = c \u00d7 e + e 2 \u03b8 = \u03b8 \u2032 \u03d5 = \u03d5 \u2032 i f ( \u03d5 \u2032 \u2265 \u03c0 ){ p = c \u00d7 e + e 2 \u03b8 = \u03b8 \u2032 i f ( \u03b8 \u2032 \u2265 \u03c0 ) else \u03b8 = \u03c0 + \u03b8 \u2032 \u03d5 = 2 \u03c0 \u2212 \u03d5 \u2032 Corollary 2",
            " Please cite this article as: J.F. Myoupo et al., Fault-tolerant and energy-efficient routing protocols for a virtual three-dimensional wireless sensor network, Computers and Electrical Engineering (2018), https://doi.org/10.1016/j.compeleceng.2018.02.012 10 J.F. Myoupo et al. / Computers and Electrical Engineering 0 0 0 (2018) 1\u201316 5.2.2. Calculation of the distance ( p ) The distance between two points A and B of the space is provided by the norm of vector \u2212\u2192 AB , denoted \u2016 \u2212\u2192 AB \u2016 or only AB . According to Fig. 5 b, an actuator that moves from the sink node (with coordinates (0, 0, 0)) to point M (with coordinates ( x, y, z )) must cover distance p = \u2016 \u2212\u2212\u2212\u2192 sinkM \u2016 = \u2016 ( x \u2212 0 , y \u2212 0 , z \u2212 0 ) \u2016 = \u2016 ( x, y, z ) \u2016 . However, \u2016 \u2212\u2212\u2212\u2192 sinkM \u2016 2 = ( sink M 1 ) 2 + ( M M 1 ) 2 with ( sink M 1 ) 2 = ( sink M x ) 2 + ( sink M y ) 2 and M M 1 = sink M z . \u2212\u2212\u2212\u2212\u2212\u2212\u2192 \u2016 sinkM \u2016 2 = ( sink M x ) 2 + ( sink M y ) 2 + ( sink M z ) 2 = x 2 + y 2 + z 2 . Thus, \u2016 \u2212\u2212\u2212\u2212\u2192 sink M \u2016 = 2 \u221a x 2 + y 2 + z 2 . Thus, we have Corollary 3 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000683_icra.2013.6631258-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000683_icra.2013.6631258-Figure7-1.png",
        "caption": "Fig. 7. 1w2 (dynamic reconfiguration manipulability measure of 2ndlink) and reconfiguration manipulability measure (RMM) [14] of 2nd-link to vertical motion.",
        "texts": [
            "8 Length, mass and coefficient of viscous friction of each link and joint are set to be 0.3[m], 1.0[kg], and 2.0[N\u00b7m\u00b7s/rad]. In this simulation, we assume that tip of link-2 and link-4 are always placed y = 0, that is, when q2 and q4 are given, q1 and q3 are set as q1 = \u2212q2/2.0 and q3 = \u2212(q2 +q4)/2.0. Figure 6(a) and (b) depict the DRME (scaled) and RME (scaled) with manipulator shapes, which indicate the peak of 2nd-link-DRMM distribution and 2nd-link reconfiguration manipulability measure (RMM) distribution shown in Fig. 7. In Fig. 7, the peak of 2nd-link RMM at q2 = 90\u25e6 and q4 = 90\u25e6. On the other hand, the peak of 2nd-link DRMM at q2 = 118\u25e6 and q4 = 141\u25e6. Difference of the results is caused from considering dynamical peculiarities or not. We discuss a biped robot whose definition is depicted in Fig. 8. Table I indicates length li [m], mass mi [kg] of links and joints\u2019 coefficient of viscous friction di [N\u00b7m\u00b7s/rad], which are decided based on [8]. Our model represents rigid whole body\u2014feet including toe, torso, arms and body\u2014 having 18 degree-of-freedom"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000492_j.ijheatmasstransfer.2012.07.078-Figure2-1.png",
        "caption": "Fig. 2. Distribution of the turbulence model test results.",
        "texts": [
            " Few of mentioned turbulence models are in addition to the models k\u2013e and k\u2013x (BSL model combines the advantages of k\u2013e and Wilcox) [5]. In order to test the turbulent models, the numerical model (a cylinder in the channel) was developed. Dimensions of the cylinder: diameter is 0.04 m, length is 0.023 m. Dimensions of the channel: width is 0.2 m, length is 0.023 m, height is 0.5 m. The air flows through the channel, flow velocity is 10 m/s. Fig. 1 shows the calculating grid made by the ANSYS CFX code. Five points on the surface of cylinder are marked in Fig. 2. Points are located clockwise. Test results are presented at Table 1. As can be seen, turbulence modeling results are fairly good and confirm to the experimental data. k\u2013e model was selected for the further calculations, as more suitable for modeling turbulent flow at the small flow velocities. where lt \u2013 the eddy viscosity or turbulent viscosity, which must be modeled using k\u2013e model, Cl is a constant equal to 0.09 and k \u2013 the turbulence kinetic energy; e is the turbulence eddy dissipation. The values of k and e come directly from the differential transport equations [5]: @\u00f0qk\u00de dt \u00fer\u00f0qUk\u00de \u00bc r l\u00fe lt rk rk \u00fe Pk \u00fe Pkb qe; \u00f02\u00de and @\u00f0qe\u00de dt \u00fer\u00f0qUe\u00de \u00bc r l\u00fe lt re re \u00fe e k Ce1 Pk \u00fe Peb\u00f0 \u00de Ce2qe\u00f0 \u00de; \u00f03\u00de where Ce1, Ce2, rk and re are constants 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003599_j.procir.2018.04.033-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003599_j.procir.2018.04.033-Figure15-1.png",
        "caption": "Fig. 15 Another defect: loss of thickness and loss of edge",
        "texts": [
            " Thus, the temperature gradient is the main physical phenomenon in the apparition of the warping defect. To tackle the warping defects on the overhang structure, support structures could be introduced. The supports are useful to dissipate the heat, to stiffen the surface and thus to limit its deformation and to anchor the surface to the starting plate or in the consolidated powder [28]. The authors have produced specimens with supports, see Fig. 14. In these conditions, the warping defect do not occur but other defects can be noticed, like the loss of thickness and / or the loss of edge, see Fig. 15. The loss of edge defect is a loss of side geometry at the end of the overhang portion, see Fig. 16. This geometric defect may be explained by the decrease of the layer length due to the cooling phenomenon from previous melted layer to the new one. Once the first layer of a surface has cooled, it shrinks due to the temperature gradient (from 1650\u00baC to 750\u00baC). The powder of the next layer is distributed and melted. An offset is produced between successive layers. The offset between two layers would be caused by the shrinkage between 1650 and 750\u00baC, as the upper layer is warmer, it is a little longer and it clings to the previous layer with a shift. The second layer also undergoes a narrowing that causes the first layer to shrink since the force exerted is greater than the elastic limit of the material. The first layer is plastically deformed. This phenomenon is repeated for each layer but the deformation becomes smaller due to the opposite deformation of the previous layers and the increase of the part stiffness. The gap between the two layers becomes smaller each time. The defect of loss of thickness, see Fig. 15, is related to the decrease of the melted material in the middle of the overhang part. This geometric defect is due to thermal phenomena between the melted layers and supports. The defect of loss of thickness occurred on another experiment that investigated the influence of the process parameters, such as current value, offset and speed function, on the geometry of the part. In Fig. 17, three specimens were manufactured with different current value. A weak current (C) gives a better geometry because it seems to avoid the problems of overheating but decreases the melting of the powder"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003128_miltechs.2017.7988729-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003128_miltechs.2017.7988729-Figure2-1.png",
        "caption": "Figure 2. Degrees of freedom in human arm: S \u2013shoulder, E \u2013 elbow, W \u2013 wrist",
        "texts": [
            " the upper arm, the lower arm and the hand that are connected by joints and together make the arm one of the most useful and complex tool of human body. The main aim joints are shoulder joint, elbow joint and wrist joint. The arm segments can move under the control of the nervous system, muscle groups and tendons. Arms are capable of moving in different directions to perform certain tasks, such as handling, holding and shooting with a gun. In performing a certain task, the human arm is considered to have seven degrees of freedom (Fig. 2). The shoulder and the wrist give pitch, yaw and roll. The elbow only allows for pitch. As a human body\u2019s part flexes, muscles are active elements producing power and movement. Conversely, during extension, muscles generate a passive tension. During movement bones and joints are passive elements controlled by the muscles. There are several published studies, in which the muscle mathematical model is mentioned. For instance, the threeelement Hill muscle model, constituted by a contractile element (CE) and two non-linear spring elements, one in series (SE) and another in parallel (PE) (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001478_s11071-018-4460-2-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001478_s11071-018-4460-2-Figure1-1.png",
        "caption": "Fig. 1 Dynamic model of the motor-gear system",
        "texts": [
            " In this study, the nonlinear dynamic characteristics of the gear-motor system are investigated by trajectory-based stability preserving dimension reduction (TSPDR) methodology. In the TSPDR methodology herein, the CCCOI-RM is chosen and the stability margins are specially defined for distinguishing the stable motion modes of the motor-gear system, to make the TSPDR methodology be used in the nonlinear analysis of the gear-motor system. At last, combined with modal analysis, the relationship between the stability and resonance of the gear-motor system is revealed. Figure 1 illustrates the dynamic model of the motorgear system. The symbols \u03b8m, \u03b81, \u03b82, and \u03b8l denote the angular displacements of the electric motor rotor, driving gear, driven gear, and load, respectively. The symbols Jm, J1, J2, and Jl denote the rotational inertias of the electric motor, driving gear, driven gear, and load, respectively. The symbols km1 and cm1 denote the connecting stiffness and damping between electric motor and driving gear, respectively. The symbols km and cm denote the meshing stiffness and damping between the driving gear and driven gear, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000249_elektro.2012.6225639-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000249_elektro.2012.6225639-Figure3-1.png",
        "caption": "Fig. 3. FEM analysis of the SRM, flux lines of th coil \"B\" turns are in short circuit, a) aligned rotor rotor position.",
        "texts": [
            "00 \u00a92012 IEEE 206 model configurations (changes of the m or other parameters) during the executio The static characteristics of SRM ha by means of FEM under four differ conditions. Namely 0%, 20%, 50%, 70% has been in the short circuit that corresp 6 winding turns respectively. The accuracy of the result depends o mesh and accuracy of the input paramet 10.083 nodes have been used. The calcu out for each individual rotor position static condition. The rotor position \u03d1 aligned \u03d1a to unaligned position \u03d1u wit each position the current was changed w range from 1 to 28 A. In the Fig. 2 t magnetic flux lines of health SRM unaligned position can be seen. In co Fig.3, there is the distribution of magn fault SRM for aligned and unaligned p coil \"B\" turns are in short circuit. Magnetic flux linkage calculation The first parameter which has been a linkage versus phase current for differen =f(I,x). The area bounded by maximal by both \u03c8-I curves for aligned and una equal to mechanical energy, which electromagnetic force [1]. In the Fig. 4 curves obtained by means of FEM f positions, if the phase current has been Amps (45o is equal to unaligned rotor p equal to aligned rotor position)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002012_msec2014-4029-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002012_msec2014-4029-Figure7-1.png",
        "caption": "FIG 7: TEMPERATURE PROFILE AFTER (a) 3 sec, (b) 12 sec",
        "texts": [],
        "surrounding_texts": [
            "To validate the finite element model, two different experimental cladding conditions have been used as mentioned in Table 5. The experimental values of dilution, heat affected zone and longitudinal residual stress are compared with those obtained from the finite element model."
        ]
    },
    {
        "image_filename": "designv11_13_0000929_b978-0-444-52215-3.00008-8-Figure8.8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000929_b978-0-444-52215-3.00008-8-Figure8.8-1.png",
        "caption": "FIGURE 8.8 Cross-sectional diagram of a shell-and-tube heat exchanger showing the thermal resistances and the modes of heat transfer. One fluid (inside yellow) flows through an inner tube that has a wall thickness (inner grey circle) and transfers energy with another fluid (outside pink) that flows through an outer tube that has a wall thickness (outer grey circle). At each boundary along radius, r, heat is transferred according to the heat transfer mode. Heat transfers through a solid by conduction (solid grey radial areas) and through solid\u2013fluid interfaces (dashed line circles) by convection. Heat transfer through the insulation is by conduction (patterned grey area). The heat exchanger can lose or gain energy by radiation or convection with the surroundings. The diagram is very detailed. Sometimes, the main interest of the analysis is the tube-side fluid. For this case, it is possible to develop an expression for the overall heat transfer coefficient in terms of the inner tube radius, r1 (\u00bcri) or the outer tube radius, r2 (\u00bcro).",
        "texts": [
            " Thus, conductive modes of heat transfer is intimately related to convective modes of heat transfer. Convective modes of heat transfer occur according to fluid flow around a solid or surface. Convective modes of heat transfer are greatly influenced by both flow velocities and fluid properties. Radiative modes of heat transfer occur anytime that there is a temperature difference between two bodies. Radiative heat transfer becomes important when the temperature differences between two surfaces are large. A cross-section of a shell-and-tube heat exchanger is shown in Figure 8.8. For this case, consider that the hottest temperature is in the center of the device and T0>T1>T2> T6, where T6 represents the temperature of the Chapter 8 Heat Transfer and Finite-Difference Methods 587 surroundings. Suppose that only purely conductive modes of heat transfer are present, then Fourier\u2019s law of heat conduction, Eq. (8.6) would apply. For the case of radial geometry, the area normal to the thermal gradient would be equal to 2prL, where r is the radius and L is the length of the cylinder",
            " The thermal resistance for convective heat transfer is: Rth, j \u00bc 1 hjAj \u00bc 1 hj 2prjL convective \u00f08:51\u00de The thermal resistance for a radiative surface can be treated in the same way as convection by using the radiative heat transfer coefficient, hr, defined by Eq. (8.13). The thermal resistance for radiative heat transfer is: Rth, j \u00bc 1 hr, jAj \u00bc 1 hr, j 2prjL \u00f08:52\u00de Conductive, convective, and radiative thermal resistances given by Eqs. (8.46), (8.51) and (8.52) can be combined as sum of thermal resistances using Eq. (8.49) to describe the rate of heat transfer that occurs according to an overall temperature gradient for various thermal resistances. If this is done for Figure 8.8, the result is: q\u00bc DTX j Rth, j \u00bc T0 T6\u00f0 \u00de 1 h1 2pr1L\u00f0 \u00de|fflfflfflfflfflffl{zfflfflfflfflfflffl} convection \u00fe ln r2=r1\u00f0 \u00de ktube 2pL\u00f0 \u00de|fflfflfflfflfflffl{zfflfflfflfflfflffl} conduction \u00fe 1 h2 2pr2L\u00f0 \u00de|fflfflfflfflfflffl{zfflfflfflfflfflffl} convection \u00fe 1 h3 2pr3L\u00f0 \u00de|fflfflfflfflfflffl{zfflfflfflfflfflffl} convection \u00fe ln r4=r3\u00f0 \u00de kshell 2pL\u00f0 \u00de|fflfflfflfflfflffl{zfflfflfflfflfflffl} conduction \u00fe ln r5=r4\u00f0 \u00de kinsul 2pL\u00f0 \u00de|fflfflfflfflfflffl{zfflfflfflfflfflffl} conduction \u00fe 1 h5\u00fehr\u00f0 \u00de 2pr5L\u00f0 \u00de|fflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflffl} convection\u00feradiation \u00f08:53\u00de Chapter 8 Heat Transfer and Finite-Difference Methods 589 where convective boundaries in Figure 8.8 are given by the dashed lines. In Figure 8.8, the bulk temperatures of fluid a and fluid b are given by Ta and Tb, respectively, and Tb is the bulk temperature of fluid b that lies between T2 and T3. Equation (8.53) can be written for any thermal resistances as: q\u00bc DT S j Rth, j \u00f08:54\u00de or in terms of the heat flux, q/A as: q A \u00bc DT AS j Rth, j \u00f08:55\u00de and then the overall local heat transfer coefficient, U, can be defined in terms of the thermal resistances as: U\u00bc 1 AS j Rth, j \u00f08:56\u00de so that the rate of heat transfer can be written as: q\u00bcUADT \u00f08:57\u00de which is valid for any geometry provided suitable definitions for the thermal resistances are used",
            "20), since the flow region is an annulus: Nu\u00bc ho LC k \u00bc ho dh k \u00bc ho k 4 Across-section Lwettedperimeter \u00bc ho k Di do\u00f0 \u00de \u00f08:59\u00de Introduction to Supercritical Fluids590 where Di and do refer to the inside diameter of the outer tube and outside diameter of the inner tube, respectively, of a shell-and-tube heat exchanger. Do the calculation of dh and check Eq. (8.59)! Example 8.6 Expression for the overall heat transfer coefficient Liquid carbon dioxide is heated by hot water in a shell-and-tube heat exchanger. Neglecting heat losses through the shell wall, shell insulation, and to the surroundings, determine an expression for the overall heat transfer coefficient based on the outside area of the heat exchanger tube. Solution. Figure 8.7 shows an image of the heat exchanger. Figure 8.8 shows the detailed heat transfer resistances and heat transfer modes that can be simplified as shown in the figure to the right. For the inner tube, use ri for the inside radius and ro for the outside radius. Use ktube for the thermal conductivity of the tube. Conductive heat transfer modes are given according to Eq. (8.46) and convective heat transfer modes given according to Eq. (8.51). Using the outside tube area, (2proL) in Eq. (8.56) gives Uo: Uo \u00bc 1 2proL 1 1 hi 2priL\u00f0 \u00de \u00fe ln ro=ri\u00f0 \u00de 2pktubeL \u00fe 1 ho 2proL\u00f0 \u00de \u00bc 1 ro rihi \u00fe ro ln ro=ri\u00f0 \u00de ktube \u00fe 1 ho \u00f0a\u00de The Uo denotes the local overall heat transfer coefficient that is a function of all fluid and material properties"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002029_10402004.2014.983250-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002029_10402004.2014.983250-Figure10-1.png",
        "caption": "Fig. 10\u2014Sliding and tearing of a surface crack based on the system of two cantilevers.",
        "texts": [
            " 9b it can be seen that mode II SIF has the same variation rule with that under distinct crack inclination angles; that is, mode II SIF initially increases from 0 to the maximum value as the roller element approaches the surface crack and then reduces to the minimum value, from which mode II SIF decays again to 0 as the roller element abandons the surface crack. However, the range of mode II SIF under various crack depths has a reverse variation rule with that under distinct crack inclination angles, as shown in Fig. 9d. The range of mode II SIF gradually increases with increasing crack depth. This fact can be explained by the shear theory based on the system of two cantilevers, as shown in Fig. 10: with increasing crack depth, the length of the crack is increased and thus the cantilevers are elongated, leading to reducing the shear area. The reduced shear area enables the sliding damage to occur when the load (Fs) remains unchanged. Therefore, the surface crack depth significantly influences the mode II RCF damage; that is, the larger the surface crack depth is, the more dangerous the mode II crack propagation is and vice versa. Figure 9c shows the variation in mode III SIF under distinct crack depths"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001988_3dp.2014.0017-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001988_3dp.2014.0017-Figure1-1.png",
        "caption": "Figure 1. The metal - based selective inhibition sintering (SIS) process.",
        "texts": [
            " In the current metal process (SIS - metal), a commercial piezoelectric printhead is utilized to deposit a liquid chemical solution (inhibitor) at the periphery of the part for each layer. Once all of the layers have been completed, the entire part is removed from the machine and bulk sintered in a conventional sintering furnace. The inhibitor deposited at the part \u2019 s boundary decomposes into hard particles that impede the sintering process. The particles in this region are prevented from fusing, allowing for removal of inhibited boundary sections and revealing of the completed part. It is easiest to think of the part as if it were encased in a sacrificial mold. Figure 1 is an illustration of the SIS - metal process. Advantages of the SIS - metal process include the following: \u2022 The machine will be orders of magnitude lower in cost due to the use of a commercial printhead instead of expensive laser or electron beam generators. \u2022 The process is faster since only the boundary of the part is treated, whereas in other methods, the entire part cross section is scanned or treated. The speed of SIS would be significantly higher if vector (instead of raster) printing is used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure18.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure18.1-1.png",
        "caption": "Fig. 18.1 Conditional distribution of the leakage fluxes and fluxes of self- and mutual induction of an induction machine with a squirrel-cage solid rotor (a) leakage fluxes and fluxes of self- and mutual induction of the stator winding; (b) leakage fluxes and fluxes of self- and mutual induction of the rotor winding; (c) leakage fluxes and fluxes of self- and mutual induction of the eddy currents induced in the rotor tooth region; (d) leakage fluxes and fluxes of self- and mutual induction of the eddy currents induced in the rotor joke region",
        "texts": [
            " In this chapter, we consider the circuit loops created by the eddy currents induced in a squirrel-cage solid rotor and their impedance values at a weak skin effect. An electric machine with a squirrel-cage solid rotor can be represented as a system of four inductively coupled windings (circuits): the stator winding (1), squirrel-cage type rotor winding (c), and circuits of the teeth (z), and rotor joke (a) region. The conditional pictures of the field distribution created by the current of each winding (circuit) individually are presented in Fig. 18.1. We use Fig. 18.1 to obtain the voltage equations for the windings (circuits) of an electric machine with a squirrelcage solid rotor. Then, using the field distribution pictures (Fig. 18.1) and the system of the equations obtained in (2.87), we have U1 \u00bc r1I1 \u00fe jx1I1 \u00fe jxc1Ic \u00fe jxz1Iz \u00fe jxa1Ia 0 \u00bc rcc s Ic \u00fe jxcIc \u00fe jx1cI1 \u00fe jxzcIz \u00fe jxacIa 0 \u00bc rcz s Iz \u00fe jxzIz \u00fe jx1zI1 \u00fe jxczIc \u00fe jxazIa 0 \u00bc rca s Ia \u00fe jxaIa \u00fe jx1aI1 \u00fe jxcaIc \u00fe jxzaIz \u00f018:1\u00de where Ic is the current and xc is the total reactance of self-induction of the squirrelcage type rotor winding, and x1c, xc1, xcz, xzc, xca, xac are the reactance values of mutual induction for the corresponding pairs of stator and rotor windings (circuits). \u00a9 Springer International Publishing Switzerland 2015 V. Asanbayev, Alternating Current Multi-Circuit Electric Machines, DOI 10.1007/978-3-319-10109-5_18 687 The magnitudes of the voltage and currents used in (18.1) are effective values. The total reactance of self-induction x1, xc, xz, xa can be represented as the sum of the reactance of self-induction and leakage reactance values. Then, we have on the basis of Fig. 18.1 that x1 \u00bc x11 \u00fe x1\u03c3 xc \u00bc xcc \u00fe xc\u03c3 xz \u00bc xzz \u00fe xz\u03c3 xa \u00bc xaa \u00fe xa\u03c3 \u00f018:2\u00de In (18.2), the values of x1\u03c3, xc\u03c3, xz\u03c3, xa\u03c3 are determined by the leakage fields created by the currents of the stator and rotor windings (circuits) (Fig. 18.1), and x11, xcc, xzz, xaa represent the reactance values of self-induction of the stator and rotor windings (circuits). On the basis of Fig. 18.1, for the leakage reactance values xc\u03c3, xz\u03c3, xa\u03c3 xc\u03c3 \u00bc xcc\u03c3 \u00fe x\u03a0c \u00fe x\u03c402 xz\u03c3 \u00bc xcz\u03c3 \u00fe x\u03a0z \u00fe x\u03c402 xa\u03c3 \u00bc xca\u03c3 \u00fe x\u03c4cz \u00fe x\u03c402 \u00f018:3\u00de In a specific system of units, the reactance values of self- and mutual induction of the stator and rotor windings (circuits) are similar, and they are equal to the magnetizing reactance xm, i.\u0435., x11 \u00bc xcc \u00bc xzz \u00bc xaa \u00bc xm x1c \u00bc xc1 \u00bc x1z \u00bc xz1 \u00bc xa1 \u00bc x1a \u00bc xm \u00f018:4\u00de The reactance of mutual induction for the corresponding pairs of rotor windings (circuits) can be represented as (Fig. 18.1) xcz \u00bc xm \u00fe x\u03c402 \u00fe x\u03a0z xzc \u00bc xm \u00fe x\u03c402 \u00fe x\u03a0c xac \u00bc xca \u00bc xza \u00bc xaz \u00bc xm \u00fe x\u03c402 \u00f018:5\u00de According to Appendices A.13.1 and A.14.1, the conditions shown in (18.3), (18.4), and (18.5) are also satisfied for the reactance values expressed in a phase system of units. In this case, the referred values are used in (18.3), (18.4), and (18.5). For the magnetizing current, we have Im \u00bc I1 \u00fe Ic \u00fe Iz \u00fe Ia \u00f018:6\u00de Using expressions (18.2), (18.3), (18.4), (18.5), and (18.6), the system of equations (18.1) obtains 18",
            " In relation to (19.8), the wound part of the rotor tooth region containing the squirrel-cage type winding is represented as a conditional planar layer with a thickness equal to hc (Fig. 19.1b). As was shown in Chap. 17, the bottom region of the rotor slot can be represented as a conditional conducting semi-infinite layer. The rotor crown region is replaced by an analogous conditional layer. As a result, the squirrel-cage solid rotor model becomes a four-layer system (Fig. 19.1b). In accordance with Fig. 18.1b, the squirrel-cage solid rotor model consists of conditional layers reflecting the air gap with a length equal to \u03b4/2, a rotor tooth region with a thickness equal to hc and slot bottom and tooth crown regions. The tooth crown conditional layer is represented as being combined with the rotor tooth layer (Fig. 19.1b). We proceed from the equivalent circuit constructed in Chap. 17 for the slotted solid rotor (Fig. 17.2). An equivalent squirrel-cage solid rotor circuit can be obtained in the form in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003666_access.2018.2879649-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003666_access.2018.2879649-Figure4-1.png",
        "caption": "FIGURE 4. Six degrees of freedom for aircraft [13].",
        "texts": [
            " In addition to studying all the systems of longitudinal motion and lateral motion of UAV, we also designed a variety of controllers for 12 different systems and improved the design of the non-minimum phase system that could not be overcome in C.-Y. Yang\u2019s thesis, and successfully developed the \u2018\u2018Reverse Gain PID Controller\u2019\u2019. Therefore, this paper is currently themost comprehensive one in the study of controller design for UAV motion. B. MODEL ESTABLISHMENT OF THE UAV The basic dynamic motion of the aircraft has six different degrees in terms of freedom, that is, the horizontal movement and the rotation around the X, Y, and Z axes, respectively (shown in FIGURE 4 [13]), which can be used to analyze the longitudinal and horizontal motions of the UAV movement. In this research, the experimental UAV, scale Cessna 182, has the scale ratio of 1: 6.65 of the prototype Cessna 182, and it has similar system performances to the prototype aircraft. In addition, because the basic dynamics and the derivation process of the aircraft model are of general knowledge, this paper only gives a brief description on how to derive the formula process of the UAV model and highlights the system analysis and design of various controllers. The detailed information for the derivation process of the UAV mathematical model can be referred in [14] or can be easily found on the internet. C. THE LONGITUDINAL MOTION EQUATIONS OF SCALE CESSNA 182 The following is a brief derivation and description of the longitudinal motion equation for aircraft. Again, the complete derivation of aircraft motion\u2019s equation can be referred in reference [14]. In FIGURE 4, as we know, the longitudinal linearization equation of the aircraft can also be represented by equation (1). u\u0307 = \u2212gcos21\u03b8 + ( Xu + XTu ) u+ X\u03b1\u03b1 + X\u03b4E\u03b4E VP1 \u03b1\u0307 = \u2212gsin2C Zuu+ Z\u03b1\u0307\u03b1\u0307 + Z\u03b1\u03b1 + ( Zq + VP1 ) \u03b8\u0307 +Z\u03b4E\u03b4E \u03b8\u0308 = ( Mu +MTu ) u+ ( M\u03b1 +MT\u03b1 ) \u03b1 +M \u03b1\u0307\u03b1\u0307 +Mq\u03b8\u0307 +M\u03b4E\u03b4E (1) Where u\u0307 is the derivative of the forward speed U, and \u03b1\u0307 is the derivative of the angle of attack (AOA), \u03b1. Besides, \u03b8\u0307 and \u03b8\u0308 are the first and second derivative of the pitch angle \u03b8 , respectively. 21 is the pitch angle in the steady state. Besides, the other parameter values are also shown in Table 2 [14]\u2013[16], Table 3 [15], [16], and the FIGURE 4. Using the Laplace transformation to equation (1) and assuming the initial conditions are zero, the initial position of the system being balanced, we can get the result as follows: L (\u03b4E) = \u03b4E (s) L(u) = u(s); L(u\u0307) = su(s) L(\u03b1) = \u03b1(s); L(\u03b1\u0307) = s\u03b1(s) L(\u03b1) = \u03b8 (s); L(\u03b8\u0307 ) = s\u03b8 (s);L(\u03b8\u0308 ) = s2\u03b8(s) When you substitute the above transformation results in Eq. (1), we can get Eq. (2):( s\u2212 (Xu + XTu ) ) u(s)\u2212 X\u03b1\u03b1 (s)+ X\u03b1\u03b1 (s)+ gcos21 (s) = X\u03b4E\u03b4E(s) \u2212Zuu (s)+ ( s ( VP1 \u2212 Z\u03b1\u0307 ) \u2212 Z\u03b1 ) \u03b1 (s) + ( \u2212s ( Zq + VP1 ) + gsin21 ) \u03b8 (s) = Z\u03b4E\u03b4E \u2212 (Mu +MTu) u(s)\u2212 (M \u03b1\u0307s+ (M\u03b1 +MT\u03b1 ))\u03b1(s)+ s(s\u2212Mq)\u03b8 (s) = M\u03b4E\u03b4E (2) 70736 VOLUME 6, 2018 Now, let the angle of the elevator, \u03b4E (t), be the system input, and the output variables of X, Y, and Z axes are u (t), \u03b1 (t), and \u03b8(t), respectively",
            " EQUATIONS OF LATERAL MOTION The linearized equation of lateral motion for scale Cessna 182 after deriving can be seen in Eq. (10) [14]:( VP1 \u03b2\u0307+VP1\u03c8\u0307 ) = g\u03c6+Y\u03b2\u03b2+Y \u03c6\u0307\u03c6\u0307+Y \u03c8\u0307 \u03c8\u0307+Y \u03b4A\u03b4A+Y \u03b4R\u03b4R \u03c6\u0308 \u2212 IXZ IXX \u03c8\u0308 =L\u03b2\u03b2+L\u03c6\u0307\u03c6\u0307+L\u03c8\u0307 \u03c8\u0307+L\u03b4A\u03b4A+L\u03b4R\u03b4R \u03c8\u0308 \u2212 IXZ IZZ \u03c6\u0308=N\u03b2\u03b2+N \u03c6\u0307\u03c6\u0307+N \u03c8\u0307 \u03c8\u0307+N\u03b4A\u03b4A+N\u03b4R\u03b4R (10) Where \u03b2\u0307 is the differential of sideslip angle, \u03b2. 9\u0307 and 9\u0308 are the first and second differential of yaw angle, 9. Likewise, \u03c6\u0307 and \u03c6\u0308 are the first and second differential of roll angle, \u03c6, respectively. The other parameter values can be seen in Table 2, Table 6 [16], and FIGURE 4. In addition, the change rate of Euler angle and rotational angular velocity of the UAV with the small interference can be seen as the same, so we can obtain the following result: Y \u03c6\u0307 = YP; Y \u03c8\u0307 = Y r; L\u03c6\u0307 = Lp; L\u03c8\u0307 = Lr; N \u03c6\u0307 = Np; N \u03c8\u0307 = Nr Then, using the Laplace transformation to Eq. (10) and assuming the initial conditions are zero; also the initial flying state of the system is stable; therefore, we can also get the following result: L (\u03b4A) = \u03b4A (s) ; L (\u03b4R) = \u03b4R (s) L (\u03b2) = \u03b2 (s) ; L ( \u03b2\u0307 ) = s\u03b2 (s) L (\u03c6) = \u03c6 (s) ; L ( \u03c6\u0307 ) = L (p) = s\u03c6 (s) ; L ( \u03c6\u0308 ) = L (p\u0307) = s2\u03c6 (s) L (\u03c8) = \u03c8 (s) ; L ( \u03c8\u0307 ) = L (r) = s\u03c8 (s) ; L ( \u03c8\u0308 ) = L (r\u0307) = s2\u03c8 (s) 70738 VOLUME 6, 2018 Set I1 = IXZ IXX , I2 = IXZ IZZ , and substitute I1 and I2 to the Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003135_0954406217720823-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003135_0954406217720823-Figure3-1.png",
        "caption": "Figure 3. The constraining forces and the deformation of the ring.",
        "texts": [
            " As observed in Figure 2, considering the initial stress caused by the uniform pressure acting on the inner wall of ring, 0 denotes the initial stress given by the following 0 A \u00bc 1 2 Z 0 A 2r\u00fe p0b sin rd \u00bc p0br\u00fe Ar 2 2 \u00f010\u00de where 0 is the initial stress, A is the ring section, is the density, is the rotational speed, b is the belt effective width, and p0 is the internal pressure. The radial, tangential constraining forces, and moments are also included to model real engineering structures, as shown in Figure 3. Energies in the Hamilton principle. All energies can be expressed as follows U\u00bc b Z 2 0 Z h 2 h 2 1 2 \" \u00fe 1 2 r \"r \u00fe 1 2 z\" z\u00fe 0 \" rd \u00f0r R\u00ded \u00fe Z 2 0 1 2 kww 2 \u00fe 1 2 kvv 2\u00fe 1 2 kuu 2\u00fe 1 2 ku bop 2 2 rd \u00f011\u00de T \u00bc Z 2 0 1 2 Ar\u00bd _ u v 2 \u00fe _ v\u00fe u\u00fe r\u00f0 \u00de 2 \u00fe _ w 2 d \u00f012\u00de W \u00bc bp0r Z 2 0 u\u00fe 1 2r v2 vu0 \u00fe v0u\u00fe u2 d \u00fe Z 2 0 quu\u00fe q v\u00fe q v0 u r rd \u00f013\u00de where h is the ring thickness, qu, qv, and q are the applied radial force, tangential force, and moment, respectively. bop is the nominal width of tire. Considering the length of the sidewall to the center line of tread is different from a half of the belt effective thickness b, the potential energy of sidewall should be corrected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000961_s10237-017-0874-x-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000961_s10237-017-0874-x-Figure13-1.png",
        "caption": "Fig. 13 Stress contour (Pa) for the terminal slip in the flexion movement at the DIP joint: a fifth-order Ogden, b invariants model, and contour of nominal c shear strain, d maximum principal strain",
        "texts": [
            " Also, the locations of the largest stress in both models showed different patterns, because as Fig. 6b, c indicates, the pure shear and balance biaxial tension modes of the Ogden model show the asymptotic behavior at lower stretches. Consequently, in stress analysis of the tissue under the flexion movement, this behavior may lead to miscalculations in prediction of stress distribution under flexion. The results of the simulation of the terminal slip part in the flexion movement at the DIP joint are presented in Fig. 13. As discussed in Sect. 3.2, the fifth-order Ogden model represents less stresses in pure shear (Figs. 6b, 7b) and balance biaxial tension (Figs. 6b, 7c) modes for terminal slip than central slip. As a result, the von Mises stress in Fig. 13a is less than Fig. 12a. According to Fig. 13c, the dominant mode of deformation is shear in the yz plane because the terminal slip behaves like a short beam. The deformation measure of Fig. 7 (horizontal axis) is stretch. But, due to shear modes in Fig. 13c, the principal values of strain tensor should be computed. Therefore, Fig. 13d shows the contour of maximum principal nominal strain. Converting the principal nominal strain to stretch using \u03bb = 1 + \u03b5, from Fig. 13d, it can be concluded the maximum stretch is 1.08906. As mentioned, the dominant mode of deformation is shear. By comparing Figs. 7b and 11b, one can observe at stretches of about 1.09 that the shear stress is less for Ogden N = 5, and this causes fewer stress contours in Fig. 13a with respect to Fig. 13b. But, as discussed, the more accurate behavior of invariants model is due to stable (non-asymptotic and non-decreasing stress) behavior. The study identified the model that best represents the hyperelastic properties of the extensor apparatus. The experimental data from different parts of the extensor apparatus were used to choose the best hyperelastic model. Different hyperelastic strain energy density functions were employed to characterize the behavior of the extensor apparatus under tension tests"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002395_58832-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002395_58832-Figure1-1.png",
        "caption": "Figure 1. Generalized coordinates of the car-like robot",
        "texts": [
            " The experimental results are presented in section 6. Section 7 concludes the paper and outlines future research. 2. Kinematic Model of the Car-like Non-holonomic Mobile Robot In this paper, the kinematic model of a car-like mobile robot is used. The main feature of the kinematic model of the car-like mobile robot is its non-holonomic constraints, because a rolling-without-slipping condition exists between the wheels and the ground. The generalized coordinates of the car-like robot model is shown in Figure 1, where the rear wheels are driving wheels and the front wheels are steering wheels. The radius of the wheels is \u03c1 . The car-like robot is round; O is the centre, where the Cartesian coordinate of O is (x, y) and the radius of the car-like robot is Ro. Point W is the midpoint of the rear wheel axis and point V is the midpoint of the front wheel axis. l is the distance between M and V. \u03b8 measures the orientation of the car body with respect to the x -axis. \u03d5 is the steering angle. Let 1u be the angular velocity of the driving wheels and 2u be the steering rate of the front wheels; thus, the kinematic model for the car-like robot can be obtained by equation (1)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002720_978-3-319-24055-8-Figure3.6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002720_978-3-319-24055-8-Figure3.6-1.png",
        "caption": "Fig. 3.6 Definition of \u201cL\u201d and \u201cT\u201d driveline layouts",
        "texts": [
            " 3 Reducing Noise in an Electric Vehicle Powertrain \u2026 37 After careful development of this method it was confirmed that this simplified method was able to give results that were satisfactorily similar to those from the analysis of the fully detailed design. Once the concept layout for the rotating components had been selected, there was still the decision as to how the driveline and power electronics were to be assembled into the vehicle. Two options were identified, referred to as \u201cT\u201d and \u201cL\u201d layouts (Fig. 3.6). Naturally, it was not possible to fully design a housing for both layouts; the project timing required that the choice had to be made without either design being fully modelled. Therefore, a simple, representative housing was modelled for both structures and the system simulation carried out. This time the simulation was more involved. The driveline mount stiffnesses were included and the assessment was made based on total structure borne vibration, measured at the mounts, and by summing the housing kinetic energy, which is indicative of the total radiated noise"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002351_s12206-015-0506-2-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002351_s12206-015-0506-2-Figure1-1.png",
        "caption": "Fig. 1. Three possible scenarios for the ground and the contact nodes.",
        "texts": [
            " Recommended by Guest Editor Sung(Soo Kim and Jin Hwan Choi \u00a9 KSME & Springer 2015 the non(rolling mesh concept allows a different derivation of the equations of motion that is based on a mixed Eulerian( Lagrangian description. One immediate outcome is that a fixed set of nodes can be chosen as contact nodes, and as a result, a CMS method can be applied to the rolling tire. It is noteworthy to mention that the selected set of contact nodes might not actually always be in contact with the ground; i.e., they are just candidates for becoming in contact with the ground. Fig. 1 shows three possible scenarios for the ground and the contact nodes. In a discretized tire model, including a high(resolution model, the total number of nodes in contact with the ground can change from one time(step to the next, even if the actual size (Length) of the contact patch remains constant. This causes an oscillatory exchange of mass between the free and the constrained portions of the tire, which will manifest itself as an artificial high frequency in the response. This critical phenomenon associated with a discretized tire model in a roll( ing scenario can be eliminated by using a non(rolling mesh [15(17]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002485_s00170-016-8430-x-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002485_s00170-016-8430-x-Figure1-1.png",
        "caption": "Fig. 1 Geometry of a hypoid gear",
        "texts": [
            " The distance between the root cone and pitch cone apices is represented as A1, whereas the addendum height at the heel side of the gear relates the face and pitch cones. Face width F is defined as the distance between toe and heel along the generatrix of the pitch cone. A point P is the mid-point of a line passing through the face width of the gear. The gear teeth cross-section is defined in a plane \u03a0, which is perpendicular to the root cone generatrix and passing through the point P. The intersection between the plane \u03a0 and the root cone generatrix forms a point N as shown in the Fig. 1. In a cross-section of hypoid gear teeth, the point N is attached to the mid-point of the land, whereas it lies at the end of an axis s!, normal to the land and coincide with the line NP ! as shown in the Fig. 2. Amean pointM lies in the\u03a0 plane at a height of hM, which is the average of gear and pinion dedendums in the \u03a0 plane. Height hM is calculated using the formulae given in [14]: hM \u00bc bG \u00fe bP\u00f0 \u00de=2 \u00f01\u00de Dedendums of the gear and pinion in the \u03a0 plane can be calculated by [19] bG \u00bc boG\u22120:5\u22c5F\u22c5sin\u03b4G \u00f02\u00de bP \u00bc aoG\u22120:5F\u22c5sin \u0393o\u2212\u0393R\u2212\u03b4G\u00f0 \u00de \u00fe c \u00f03\u00de Where aoG and boG are the addendum and dedendum of the gear at the heel, \u03b4G is the dedendum angle and c is the clearance between the gear and the pinion teeth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000367_cdc.2013.6760583-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000367_cdc.2013.6760583-Figure5-1.png",
        "caption": "Fig. 5. Example 2 - CSC path with \u03b80 = 0 and \u03b8f = \u03c0 12",
        "texts": [
            " 3) C3 type of curve if A = 0, 1 \u03b63 (s) = 1 \u03b60 \u2212Bs 4) C4 type of curve if B = 0, \u03b64 (s) = \u03b60 +As \u03b8, x and z can be deduced from the expression of \u03b6 as follows:          \u03b8 (\u03b6) = 2 arctan \u03b6 x (\u03b6) = x0 + zr (\u03b8 (\u03b6)\u2212 \u03b80)\u2212 zr (A+B) s (\u03b6) z (\u03b6) = z0 \u2212 zr ln ( 1 + \u03b620 A+B\u03b620 A+B\u03b62 1 + \u03b62 ) Once again, Theorem 3.4 applies. Therefore, optimal paths can be of type CSC and CSCSC, C denoting one of the four types of arcs presented above. For this application case, a reflection can be renamed a rebound. Figures 5, 6 and 7 present some paths with line segments that are candidates for optimality. The maximum curvature at altitude zero is chosen as c(0) = 0.0015m\u22121, it depends on the glider parameters. Figure 5 presents a C1SC1 path starting close to the ground with initial orientation angle \u03b80 = 0 and ending at an altitude greater than 20km with final orientation angle \u03b8f = \u03c0/12. Figure 6 presents a C2SC2 path starting and ending at high altitude with initial orientation angle \u03b80 = \u2212\u03c0/3 and final orientation angle \u03b8f = \u03c0/3. Figure 7 presents two CSCSC paths starting and ending at high altitude. This paper presented the problem of minimizing path length in an heterogeneous environment for a Dubins\u2019 vehicle moving forward and steering with a maximum curvature that depends on the vehicle position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000929_b978-0-444-52215-3.00008-8-Figure8.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000929_b978-0-444-52215-3.00008-8-Figure8.1-1.png",
        "caption": "FIGURE 8.1 Modes of heat transfer: (i) conduction by boundary\u2013boundary contact, (ii) free convection between a surface boundary and the surroundings (stagnant air); (free convection is also called natural convection) (iii) forced convection between a moving boundary (air) and a fixed-surface boundary and (iv) radiation between a source at high temperature and a surface at low temperature. Forced convection and conduction are the main heat transfer modes for heat exchange between fluids.",
        "texts": [
            " The science of heat transfer uses thermodynamics to define fundamental quantities such as temperature, heat, enthalpy and work, along with basic relationships that have been hypothesized in science to describe the three modes of energy transport known as conduction, convection, and radiation. Some of the ways that heat transfer theory can be used are: (i) design and safety of heat exchange systems for efficient use of energy (ii) analysis of steady-state or unsteady-state conditions in energy storage or production (iii) prediction of property profiles within a flowing fluid to explain a reaction. Some physical examples of heat transfer that occur by conduction, convection and radiation modes are shown in Figure 8.1. Conduction is heat transfer that occurs within a body due to thermal gradients. Convection is heat transfer that occurs due to the motion of a fluid between a hot and a cold body. Free convection or natural convection is heat transfer that occurs due to fluid -0-444-52215-3.00008-8 557 movement caused by buoyancy forces. Forced convection is heat transfer that occurs due to fluid flow. Radiation is heat transfer that occurs due to electromagnetic waves. Some examples of heat transfer by the three different modes that can be seen in Figure 8.1 are: (i) Conduction. A hot block is place on a cold block (Figure 8.1). Heat transfer occurs by conduction when a hot block that is placed on a cold block so that heat is transferred by physical contact of the surfaces of the objects rather than by fluid transport. (ii) Convection. The surrounding air temperature is different from that of the hot block and the cold block (Figure 8.1). In free convection or natural convection, heat transfer occurs when air moves along a hot (or cold) surface naturally due to buoyancy effects. In forced convection, heat transfer occurs when air moves along a hot or cold surface due being blown by a fan across a hot (or cold) object. (iii) Radiation. Heat transfer occurs by radiation when electromagnetic waves are transmitted from a hot object to a cold object. In Figure 8.1, rays from the Sun warm the left surface of the cold block and also the left surface of the hot block. In the analysis of heat transfer, the transport of energy occurs by all three modes regardless of the situation. However, some modes are more important than others for a given analysis. In Figure 8.1, suppose that a hot block is placed in close contact with a cold block. A large temperature gradient will exist between the blocks. The temperature gradient will cause energy to be transported from the hot block Chapter 8 Heat Transfer and Finite-Difference Methods 559 to the cold block by conduction. The red overhanging part of the hot block will also radiate energy to the right surface of the cold block. The energy transport that occurs via radiation from the hot block to the cold block depends on the temperature difference between the two blocks and the exposed surfaces that can see each other or view factor. In the analysis of the heat transfer to the cold block in Figure 8.1, there are several sources of energy: (i) the source of the hot block that will transport energy by conduction and radiation modes (ii) the source of the sun, whose rays will hit on the exposed surfaces of the cold block that will transport energy by radiation and (iii) the source of the surrounding air that will transport energy away from the cold block by free convection mode. One can also analyze the heat losses of the hot block. Sometimes the energy transport is transient or time-dependent such as in the example of Figure 8.1 in which a hot block is placed on a cold block. Other times, energy transport can be steady-state in which temperature gradients between two bodies are maintained by constant supply and constant removal of heat. The focus of this chapter is on steady-state energy transport. In this chapter, the device introduced for transferring energy from one fluid to another fluid is the heat exchanger (see Figure 8.2). The main heat transfer modes of a heat exchanger are forced convection, due to movement of two fluids exchanging heat, and conduction, due to physical barriers that are used in the heat exchanger to separate the two fluids at different temperatures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000890_tmag.2014.2323305-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000890_tmag.2014.2323305-Figure1-1.png",
        "caption": "Fig. 1. Structure of surface mounted type VFPM machines.",
        "texts": [
            " The magnetic field and induced voltage obtained by the analytical method are compared with those obtained by the FE analysis and experiment, which validates the analysis methods presented in this paper. Manuscript received March 7, 2014; revised April 25, 2014; accepted May 4, 2014. Date of current version November 18, 2014. Corresponding author: S.-H. Lee (e-mail: shlee07@kitech.re.kr). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TMAG.2014.2323305 The structure of surface-mounted type VFPM machines is shown in Fig. 1, in which 1/3 of the analysis model comprised 14 NdFeB magnets and two AlNiCo magnets. Because of the very low coercive force, the AlNiCo magnets can be easily magnetized and demagnetized via a short pulse of positive and negative d-axis current. This results in a change in the total flux per pole. Therefore, it is possible to extend the flux weakening in the high-speed operating region and to improve efficiency. Fig. 2 shows our simplified analytical model for predicting the magnetic field, where \u03b1, \u03b2, \u03b3 , and Y the radii of each region from the zero-point coordinate of the VFPM machine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001902_pc.2015.7169961-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001902_pc.2015.7169961-Figure1-1.png",
        "caption": "Fig. 1. V22 Osprey \u2013 principial scheme of construction",
        "texts": [
            " Typical control problems one needs to consider are related not only to control the vehicle\u2019s flight itself but also to control the process of its take off, landing, and stabilization during the flight. These problems significantly depend on the flying vehicle under consideration, and the corresponding control actions need to be chosen also with respect to the size and type of the vehicle. In this paper the respective vehicle we consider is the socalled bicopter, which can be described as a two-rotor flying machine without variable tilting propellers attached to it, see Fig. 1. Note that the rotors and the body of the bicopter are not firmly connected, and the whole construction is more pendulum-like. That is, the rotors can rotate around the common axis perpendicular to the body of the bicopter. This makes bicopter one of the least stable aircraft construction. The size of the machine considered here is relatively small (see Fig. 9), which creates high demands for the speed of its control. The problems we focus on here are related mainly to the construction, signal processing, stabilization, and control of the bicopter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003705_ecce.2018.8558256-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003705_ecce.2018.8558256-Figure11-1.png",
        "caption": "Fig. 11. IM locked-rotor time-harmonic FEAs are carried out to compute the cage resistance as a function of the rotor slip frequency.",
        "texts": [
            " In this operating condition, the induced current irq can be closed to the rated one, but the flux current is much lower than the one at the rated frequency and voltage. Thus, the rotor slip angular frequency has to increase significantly, with respect the rated one. A first relation between the rotor resistance and frequency is achieved performing a set of locked-rotor simulations, increasing the frequency. The cage resistance behavior, as a function of the slip angular frequency, is shown in Fig. 11. Considering magnetostatic on-load FEA, in each working point dq stator and rotor currents are imposed as field sources in the simulation. From the field solution, the flux linkage \u03bbrd is achieved and the rotor resistance can be expressed as a function of the slip angular frequency, using the rotor voltage equation in (13): Rr = \u2212\u03c9sl \u03bbrd irq (28) that is a straight line in the plane Rr-\u03c9sl. Further, (28) is a second relationship between the rotor resistance and slip frequency, besides the locked-rotor characteristic"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001777_978-1-4419-8420-3-Figure2.11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001777_978-1-4419-8420-3-Figure2.11-1.png",
        "caption": "Fig. 2.11 Inspection process on a deformed pipeline. a Upper simulated model. Lower regions guarded by different guard points encoded in different colors. b Deformed pipe. c Deformations are detected, whose regions are extracted (and refined) and shown in red",
        "texts": [
            "10; this pipe is guarded by 12 points (a). In addition, the damaged region of the pipe is big and with complex topology (b). In this case, the robot should check from more than one guarding points in order to detect the entire shape of such a big hole. The entire defect geometry is extracted by composing boundaries detected from different guarding sites. The merged boundary loop is illustrated in (c). 2.5 Result of Simulated Experiments 33 Small deformation such as bending and erosion can also be detected in our system as shown in Fig. 2.11. The detected deformed region is colored in red. Clogging also changes the scanned geometry of the pipeline and can be detected. Figure 2.12 shows an example. The clogged solid (green) is detected, and its boundary geometry is reconstructed using height maps as illustrated in dark red. The robot will report the clogs when it is detected. In this example, the reconstruction merges the geometry of the blocking stuff from two aspects (from two guarding points) using their corresponding height maps"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000518_1.4028062-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000518_1.4028062-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of HVD model for the soft-start",
        "texts": [
            " The boundary conditions of thermal equation are: (1) initial temperature; (2) heat flux; and (3) heat convection. The initial temperature and heat convection can be given by means of experimental stand tests or the reference. So the only variable that is closely related to the research model is the heat flux which is expected to be solved with the following three models including frictional torque model, pressure model and heat flux model. 2.1 Frictional Torque Model. Consider the following model for oil flow in HVD. As shown in Fig. 2, the two identical annular disks with inner radius a and outer radius b are separated by a distance h\u00f0t\u00de in the direction of the rotation axis z, x1 is the angular velocity of the driving disk, and x2 is the angular velocity of the driven disk. The rotation of the upper disk sets the fluid in the gap in motion due to the nonslip condition at the surface of the disks. Typically, the gap between the disks is very small for HVD, so the boundary layer is fully developed and, under steady-state, axisymmetric flow conditions, the Navier\u2013Stokes equations are considerably simplified into the momentum equations using the cylindrical coordinates"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000331_s10853-013-7883-7-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000331_s10853-013-7883-7-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram showing the DMD system",
        "texts": [
            " The powder particles were mostly spherical with an average particle size of 67 lm, particle size ranging from 45 to 90 lm, and *93 % of the particles lie within the 45\u201375 lm range. Approximately 100 particles per micrograph were selected for measurements and always the largest diameter and the diameter in the direction perpendicular to the long axis were measured. The average powder porosity measured from particle cross sections was found to be approximately 0.2 %. DMD systems and specimen production A schematic of the DMD system is shown in Fig. 2. Primary components of DMD system consists of a laser generation system, a powder delivery system, a feedback control system, and a CNC motion stage [15]. Specimens were prepared at Focus: HOPE using POM DMD 505 Table 1 Nominal composition of the MetcoCu\u201338Ni alloy powder Element Cu Ni Fe and minor alloying elements wt% 61.9 37.9 \\0.2 machine (developed by POM Group) equipped with a 5 kW CO2 laser system and a POM DMD 105D machine (developed by POM Group) equipped with a 1 kW fiber coupled diode laser system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003298_978-3-319-70987-1_20-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003298_978-3-319-70987-1_20-Figure4-1.png",
        "caption": "Fig. 4. The change grad s as the coupling region degenerates",
        "texts": [
            " In all cases, the defective zones corresponded to the expected boundaries for the separation of the slip and bond regions. In [12, 13], similar damage occurred under the action of a purely normal load, which is in good agreement with the experiments discussed above. It is known that sliding of conjugate surfaces even in micro volumes begins only after having passed the stage of elastic and elastic-plastic displacement. Then grad s will in many respects be determined by the position of the interfaces of the zones. Let the region E0 be sufficiently large at the initial position, as shown in Fig. 4. Let us consider the change in the tangential stresses in the degeneracy of the cohesion region (Fig. 5). In position 1, the interfaces E0 and E+ correspond to zones of contact with relatively small specific loads (both normal and tangential). The constriction of E0 leads not only to a certain increase in TR, but also to the displacement of the interfaces into the region of increased pressures, determining the increase of grad s-position 2. However, with a further increase in the velocity v, the region of adhesion, degenerating, again exits its boundaries to the periphery of the contact spot, and the effect of concentrating shearing stresses is reduced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000513_smc.2014.6974513-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000513_smc.2014.6974513-Figure5-1.png",
        "caption": "Figure 5 The relationship between the arm angles",
        "texts": [
            " Dual Arm-angle z\u03c8 and x\u03c8 According to Figure 3 and the previous definitions, the projection of the vector e on the vector w is given by \u02c6 \u02c6 \u02c6, T wd = w(w e) w = w (12) The unit vector in the arm plane which is vertical to the vector w is given by ( )\u02c6 \u02c6 \u02c6,= = \u2212 = \u2212 Tpp p e d I ww e p (13) The unit vector in the reference plane which is vertical to the vector w is given by ( ) ( )\u02c6 \u00d7 \u00d7= = \u00d7 \u00d7 = \u00d7 \u00d7 =, w V w w w V w w Vnnn n (14) Considering (13) and (14), and the properties of dot product and cross product, we can have ( ) ( ) \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 \u02c6 c s \u03c8 \u03c8 \u00d7 = = \u00d7 = T T T n p w n p w n p (15) Therefore, ( ) ( )( )\u02c6atan2 , atan2 ,s c\u03c8 \u03c8\u03c8 \u00d7= = T Tw V p V p (16) To get dual arm-angle, we establish the same attitude axes with the base frame at the point S shown in Figure 5. The z0axis and the x0-axis are selected to form the two reference planes, denoted as S_z and S_x respectively. If z0-axis is chosen as the reference axis, we let [ ]T0 0 0 0 1=V = a , the corresponding arm angle can be calculated from (16) and denoted by z\u03c8 . On the other hand, when x0-axis is chosen as the reference axis, we let [ ]T0 0 1 0 0=V = n , the corresponding arm angle is denoted by x\u03c8 , dual arm-angle are shown in Figure 5. B. The Relationship of z\u03c8 and x\u03c8 As shown in Figure 5, the two vectors _ xn and _ zn , respectively in the reference plane S_x and S_z, are both perpendicular to the line SW. The angle between the two reference planes is denoted as \u03d5 . It is a constant value for a given end-effector pose. Taking the plane S_z as the reference plane, an arm angle, denoted as z\u03c8 , is the defined as the angle between it and the arm plane. Similarly, another arm angle x\u03c8 is determined. The two arm angles satisfy a constraint condition. According to geometry shown as Figure 5, there exists the following relationship: x z\u03c8 \u03c8 \u03d5= + (17) For one of the arm angles, such as z\u03c8 , we can use the following scalar to judge whether the algorithm singularity happens: sk = \u00d7w V (18) Where, ks is the parameter to determine the algorithm singularity and called singularity parameter. 0sk \u2248 shows that the algorithm singularity happens when using single arm angle. C. Analytical Inverse Kinematics Resolution The details to determine the reference vector and arm angle are shown in Figure 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000055_1.4027750-Figure17-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000055_1.4027750-Figure17-1.png",
        "caption": "Fig. 17 The sizes and locations of multiple inhomogeneities",
        "texts": [
            "1 GHz and i5 central processing unit (CPU) was used to calculate these cases. The case D\u00bc 0 corresponds to the lubrication problem without inhomogeneity. The computation time for D\u00bc 0 is obviously shorter than the cases considering an inhomogeneity. The computation time slightly increases with the size of the inhomogeneity as the calculation of eigenstresses and eigenstrains are only in the domain of inhomogeneity and large inhomogeneity contains more cuboidal units. 4.2 Multiple Inhomogeneities. In this part, multiple inhomogeneities are considered. As shown in Fig. 17, nine inhomogeneities are located under surface with depth h*. These inhomogeneities have the same size cx\u00bc cy\u00bc cz\u00bcD and interval distance d. The influences on the pressure and film thickness for different depths are shown in Fig. 18. As shown in Fig. 18, significant pressure fluctuation is induced by multiple inhomogeneities when they are located near the surface. Because the distance between inhomogeneities are larger than their sizes, the interactions of inhomogeneities are not notable and the pressure peak can be seen distinctly just like the superposition of the effect of single inhomogeneity"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002435_iros.2015.7353998-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002435_iros.2015.7353998-Figure1-1.png",
        "caption": "Fig. 1. General bicycle posture coordinates",
        "texts": [
            " Experimental validation of the extended model is reported for different needle-tissue combinations. This paper is organized as follow. Section II presents a brief review of the bicycle model and its 3-DOF formulation for needle steering. In Section III the proposed extended bicycle model is presented and modifications made to the bicycle model are presented. In Section IV, the proposed model is experimentally validated and the results are compared to the bicycle model\u2019s results. The kinematics of a bicycle moving in the y \u2212 z plane is shown in Figure 1. The fixed frame and the body moving frame are denoted by {A} and {B}, respectively. The origin of the moving frame is attached to the body at point P , somewhere between the two wheels. The posture of the body with respect to the fixed frame {A} is described by the Cartesian position of point P and the rotation angle \u03b8 of frame {B}. The position and orientation of the wheels in the moving frame are characterized by `i(i = f, b) being the distances between the front and back wheels and point P , and \u03b2i(i = f, b) being the orientations of the wheels with respect to body frame {B}. For the general bicycle shown 978-1-4799-9994-1/15/$31.00 \u00a92015 IEEE 4375 in Figure 1, the velocity components of wheel centers along the z and y axis can be expressed in frame {B} as Bvzi = cos(\u03b8 + \u03b2i)z\u0307 + sin(\u03b8 + \u03b2i)y\u0307 + `isin\u03b2i\u03b8\u0307 (1) Bvyi = sin(\u03b8 + \u03b2i)z\u0307 \u2212 cos(\u03b8 + \u03b2i)y\u0307 \u2212 `icos\u03b2i\u03b8\u0307 (2) in which i = f, b and [ y z \u03b8 ]T and [ y\u0307 z\u0307 \u03b8\u0307 ]T represent the 3\u00d71 posture vector of the bicycle and its time derivative in the fixed frame, respectively. The typical wheels used in a bicycle satisfy the pure rolling and non-slipping constraints. Rolling occurs on the wheel plane, thus Bvzi \u2212 \u03c9wi = 0 (3) Also, there is no slipping orthogonal to wheel plane, thus Bvyi = 0 (4) In (3), \u03c9wi is the rolling speed of the wheels"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001551_s106836661804013x-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001551_s106836661804013x-Figure3-1.png",
        "caption": "Fig. 3. Kinematic scheme of a spiral tribometer. (1) rotating disk; (2) rotation trajectory of a ball; (3) guide disk; (4) force sensor; (5) scrab; (6) spiral orbit; (7) fixed disk; (8) ball.",
        "texts": [
            " In some cases, when using a thrust bearing for creating an additional load, one of the rings is installed with an eccentricity. Among vacuum model test benches, the following ones are widely used: \u2014four-ball friction machine; \u2014a \u201cball\u2013disk\u201d test bench; and \u2014spiral tribometer (ST). The key difference of a spiral tribometer from the first two consists in the fact that lubricating material is tested in conditions of rolling friction [3\u201310]. The principle of work of ST is based on the rotation of a ball along a spiral trajectory between a movable and a fixed disk (Fig. 3). In the preset section of the trajectory, a ball makes contact with the guide vertical surface of a force sensor. The value of the force measured by the sensor is proportional to the friction force. The greater part of the ball\u2019s path is pure rolling; however, at the moment of contact of the ball with the vertical surface of the force sensor, turning and slipping occur along with rolling. This section of the friction path, which is the most tribologically loaded, is known as scrab in the English-language literature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000056_j.precisioneng.2014.11.008-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000056_j.precisioneng.2014.11.008-Figure10-1.png",
        "caption": "Fig. 10. Schematic view of the test system.",
        "texts": [
            " 9 shows the designed test stand. The prototype was set up n the table of a machining center, and the rotational speed was rovided by machine tool spindle. The spindle did not have an djustable preload. The bearing system, including bearing, shaft, ousing, and the variable preload system, was fixed on a cutting force dynamometer (Kistler model 9257B) which was in turn fixed to the machine table. One end of the shaft was fit in the bearing while its other end was attached to the spindle tool holder. Fig. 10 depicts a schematic view of the system and the forces acting within the system. As shown in Fig. 10, the preload force induced by the spring passes through the bearing to the housing and is eventually measured by the dynamometer. The dynamometer measures the axial force in the system. preload variations. In each step of the testing process, enough time w b s i r t w m p t s p r s i p i a m s s t i i m a t d t c l s r F w f p the initial preload of the system as well as the augmenting effects as given to the system in order that the temperature field in the earing system reached steady state. It is noted that the expanion of the system components due to temperature rise results in ncreasing the preload"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001042_0959651817698350-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001042_0959651817698350-Figure5-1.png",
        "caption": "Figure 5. Schematic representation for the determination of circular curvature.",
        "texts": [
            " Development of control strategy for curveconstrained trajectory In this proposed control strategy, a circular curve is used to exploit redundancy based on the concept of whole-arm manipulation.3 However, it is not restricted to circular curve only, and it can also be valid for others, that is, cubic polynomial, Bezier, B-spline curve and so on. As per proposed curve, the backbone of hyper-redundant manipulator should fall onto an arc of a circle. Schematic representation for the determination of circular curvature is shown in Figure 5. The general form of a circle equation can be written as (xi xc) 2 + (yi yc) 2 = a2 \u00f010\u00de where xi and yi are the link end coordinates, respectively; xc and yc are the center coordinates and a is the radius of circle. The purpose of this work is to fit each link ends onto a circular arc (equation 10) and then perform the motion planning of the arc. For this, three control points are needed for circular curve calculation. Hence, xc, yc and a need to be expressed in terms of the link end coordinates as x1 = rcf, y1 = rsf, x2 =x1 + l1cf1, y2 = y1 + l1sf1, x3 = x2 + l2cf12 and y3 = y2 + l2sf12",
            " With the help of these equations, center coordinates (xc, yc) and radius a can be expressed as a function of (x1, y1, x2, y2, x3, y3), and these are as follows xc = a1=2a2, yc = b1=2b2 \u00f011\u00de a= ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi (x1 xc) 2 + (y1 yc) 2 q = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi (x1 a1=2a2) 2 + (y1 b1=2b2) 2 q \u00f012\u00de where a1, a2, b1 and b2 are variables found during derivation of center coordinates (xc and yc) and radius a. With the help of coordinates of the center and radius of the circle, for simplicity, one can find remaining joint angles, that is, u3, u4, u5 and u6 from Figure 5 and its corresponding tip coordinates of links 3, 4, 5 and 6. The tip coordinates can also be evaluated analytically by assuming that all links are of equal length and substituting u2= u3= u4= u5= u6. In a conventional robot, all joints must be actuated to reach around the obstacle. The hyper-redundant manipulator can achieve this by actuating a limited number of joints. Expanding this concept to 6-DOF planar space robot, one can see the advantage of hyperredundant manipulator in the next section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001026_j.mechmachtheory.2017.01.005-Figure12-1.png",
        "caption": "Fig. 12. Von Mises with rigid housing for F r = 6 kN and \u03c9 = 500 rad/s at (a) t = 1 \u00d7 10 \u22123 s (b) t = 2 \u00d7 10 \u22123 s (c) t = 3 \u00d7 10 \u22123 s.",
        "texts": [
            " The half-clamping of the outer ring and the shaft velocity lead to a dynamical load distribution where the maximum values remained close to 2 kN and the \u201cw\u201d shape seems conserved as observed in static, Fig. 8 (b). The Von Mises stress fields with respect to time are given in supplementary material (videos) for rigid and elastic housing using the same stress computation method. Each frame is averaged over a frequency of 50 kHz with the same triangular mesh. In multimedia component 1 given in Appendix A or considering snapshots Fig. 12 , the Von Mises stress field in dynamic mode recalls the profiles of the load distribution observed in Fig. 11 (a). The maximum stress stays close to \u03c8 = 0 o and follows the contact displacement. Mechanical forces at the discretized interface are transmitted within rings resulting in noisy sollicitations. The stress magnitude is determined between 200 MPa and 300 MPa according to time. Due to the rotation, the outer ring feels dynamic loading. In multimedia component 2 given in Appendix A or snapshots Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001022_tia.2017.2683439-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001022_tia.2017.2683439-Figure5-1.png",
        "caption": "Fig. 5. The four line-start SynRMs (all are 4 poles, 400 V, 50 Hz)",
        "texts": [
            " width of 15 mm each), and these strips was placed and tested in the SST system. Using the methodology described in [1], the width of the damaged area and its magnetic properties have been estimated: this material has been defined the totally \u2018damaged\u2019 material; please consider that in the proposed approach the damaged zones are assumed as a homogeneous region. Both \u2018green\u2019 and \u2018damaged\u2019 materials were used in FEM models. III. SYRMS AND TEST DESCRIPTION The four line-start synchronous reluctance motors shown in Fig. 5 have been built in order to quantify the impact of the stator and rotor geometry on their performance, by direct or comparative measurements. 0093-9994 (c) 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. In the same figure, the alphanumeric codes used to identify the machines are reported (letter: N  non annealed, A  annealed; number: 1  120 W, 5  550 W)",
            " In order to make possible the analyses, the squirrel cages used for the motor self-starting has been removed during the rotor construction. Since the assembly could involve additional negative phenomena, the same technological processes were adopted for the motor production. The machines are equipped with four PT100 thermal sensors positioned in the stator winding (two in the slots and 2 inside the end windings). However, the prototypes cannot be tested with the rated loads and in steady state thermal conditions; for this reason, the fans have been removed by their shaft in order to minimize the windage losses (see Fig.5). To comparatively discriminate the annealing effects on the core losses, magnetizing behaviors and efficiency, the two sets of twin machines have been tested in three different ways: (i) No-load tests under sinusoidal voltage supply (18kVA, 360AMXT Pacific Power Source): the tests have been performed changing the supply frequency from 20 Hz up to 50 Hz. Imposing the nominal airgap flux, the two SyRMs were able to start at 20 Hz, but the synchronization at higher speeds required to increase the V/f ratio over its nominal value"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000249_elektro.2012.6225639-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000249_elektro.2012.6225639-Figure1-1.png",
        "caption": "Fig. 1. The investigated SRM, a) cross-section area, photo of open motor.",
        "texts": [
            " The manufacturing costs of this machine are lower in comparison with others electrical machines. These features of SRM are very important in competition with other electrical machines [1]. The SRM construction is very simple. The both stator and rotor have salient poles and only stator carries winding coils, which are suitable connected to create phase. The magnetic flux is provided by phase current to develop a reluctance torque [2], [3], [4]. The cross-section area of the three phase 6/4 SRM is shown in the Fig. 1a. In the Fig. 1b, there is photo of the investigated SRM. During industrial processes, the drive reliability is very important task from point of view fault operation. Several fault conditions of electrical drive can occurred during its operation. It could be mechanical, magnetic or electric fault of the motor. This paper is focused into electrical faults of the SRM and their impact on motor behavior. The electrical faults could be: short circuit in one coil of a phase (all turns or some turns), a whole coil is bridged by a short circuit, the whole phase is short circuited, open circuit in one coil of a phase, a short circuit between two different phases, a short circuit from one winding to ground [5]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure2.4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure2.4-1.png",
        "caption": "Fig. 2.4 Conditional distribution of the leakage fluxes and fluxes of self- and mutual induction of the induction machine windings (a) leakage fluxes and fluxes of self- and mutual induction of the stator winding; (b) leakage fluxes and fluxes of self- and mutual induction of the rotor winding; (c) leakage fluxes and fluxes of self- and mutual induction of the stator and rotor windings",
        "texts": [
            "24), for the emf equations of the stator and rotor windings, we can receive 22 2 Fundamentals of the Field Decomposition Principle E1\u0440 \u00bc jx1\u03c3I1 \u00fe E1m\u00f0 \u00de \u00bc jx1\u03c3I1 \u00fe jx11I1 \u00fe jx21I2 \u00bc jx1I1 \u00fe jx21I2 E2\u0440 \u00bc jx2\u03c3I2 \u00fe E2m\u00f0 \u00de \u00bc jx2\u03c3I2 \u00fe jx22I2 \u00fe jx12I1 \u00bc jx2I2 \u00fe jx12I1 \u00f02:26\u00de On the basis of the system of equations (2.26), it follows that for the total reactance values of self-induction of the stator and rotor windings, x1 \u00bc x1\u03c3 \u00fe x11 x2 \u00bc x2\u03c3 \u00fe x22 \u00f02:27\u00de The equations (2.26) reflect the presence of the inductive coupling between the stator and rotor windings. The equations (2.26) can be obtained by using the flux (field) distribution pictures created by the current of each winding individually (Fig. 2.4a, b). In Fig. 2.4a, b, the reactance values x11, x12, x22 and x21 represent the fluxes\u03a611\u00bc\u03a612 and\u03a622\u00bc\u03a621, and reactance values x1\u03c3 and x2\u03c3 reflect the leakage fluxes \u03a61\u03c3 and \u03a62\u03c3, respectively. The picture representing the resulting flux (field) distribution in an electric machine follows as a result of superposition of the fluxes (fields) produced by the current of each winding individually (Fig. 2.4a, b). Such field picture is shown in Fig. 2.4c, where the corresponding reactance values are used instead of the fluxes. Now, to obtain equations in the form of (2.26), first, the flux (field) distribution pictures created by the current of eachwinding individually can be built (Fig. 2.4a, b). a b c 2.1 An Induction Machine with a Single-Winding Rotor: Voltage Equations. . . 23 Then, by using these flux (field) pictures, the emf equations can be written. This technique facilitates the process of determining the emf equations for the electric machine windings. In this case, a high level of visibility and clarity of the process of setting up the emf equations is achieved. Such an approach greatly simplifies the procedure for obtaining the emf equations for the electric machine windings, particularly for deriving the emf equations for electric machines representing a multi-circuit electromagnetic system",
            " The leakage fields take place when the stator and rotor windings are 0 E1p E2 ' p I2 ' x2\u03c3'x1\u03c3 0 xm Em Im I1 \u00b7 \u00b7 Fig. 2.5 Equivalent circuit constructed with respect to the magnetizing emf Em 28 2 Fundamentals of the Field Decomposition Principle flowed by the load components of the currents (this provision will be demonstrated below in more detail). The picture of the field distribution corresponding to the operation conditions is shown in Fig. 2.6. The picture of the resulting field distribution corresponding to the system of equations (2.26) is given in Fig. 2.4c. In Fig. 2.6, the referred values of the rotor winding currents are used, and in Fig. 2.4c, the real values of the currents are used. On the basis of Figs. 2.4c and 2.6, we can conclude that the resulting field picture arising from the system of equations (2.36) using the referred secondary values has a simpler structure. In accordance with (2.26), introduction of leakage reactance value is based on the use of the provision that the leakage fields (fluxes) cover turns of only one or the other of the windings. However, this provision is not sufficient for quantitative determination of the leakage reactance values, as it does not reflect the real picture of the coverage of the stator and rotor windings by the leakage fields"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure8-1.png",
        "caption": "Fig. 8 . Stewart platform and associated notation.",
        "texts": [
            "25 s when considering collisions in the 3 R RR example of Table 1 ). Although these times may still be too large for real-time control (they can always be reduced by decreasing N j s ), they may allow for collision control algorithms during online path planning. However, the detailed analysis of this application is beyond the scope of the present paper. In this section, we will use the proposed method to analyze the workspace of the Stewart platform, considering that the degree of redundancy is r = 1 . Consider a Stewart platform as shown in Fig. 8 , used for machining (assume that the Z axis is a tool, e.g. a drill). In this robot, six linear actuators or legs of type U P S are used to control the position and orientation of a frame attached to the mobile platform. The position and orientation of frame with respect to the base frame W is defined by the ( x, y, z ) coordinates of point P of the mobile platform and by the XYZ Euler angles ( \u03b1, \u03b2 , \u03b3 ). Note tha t, in machining applications, the rotation \u03b3 of the mobile platform about the tool axis (which coincides with the Z axis of frame ) is not relevant, since the tool is continuously rotating about its own axis. Therefore, if \u03b3 is not relevant, then we can consider the robot of Fig. 8 as a redundant robot in which 6 linear actuators are used to control 5 task variables: the position ( x, y, z ) and angles ( \u03b1, \u03b2). The degree of redundancy is therefore r = 1 , so self-motion manifolds are curves in the six-dimensional joint space (which is the space of lengths \u03b8 j of the linear actuators, j = 1 , . . . , 6 ). In this example, the task space is 5-dimensional. Although Algorithm 5 can be applied in principle to obtain the barriers in task spaces with any dimension m , since the number of calculations of this algorithm grows exponentially with the dimension of the task space, the computation times would become prohibitive for m = 5 ",
            " , \u03b86 ] T ( n = 6 ). The task variables are t = [ y, z] T ( m = 2 ). The only passive variable in this example is \u03c8 = [ \u03b3 ] T ( r = 1 ). Note that, since ( x, \u03b1, \u03b2) have been fixed in order to analyze a 2-dimensional slice of the task space, these three variables can be considered as geometric parameters of the robot in our formulation. For each linear actuator j \u2208 { 1 , . . . , 6 } , we impose the condition that its length must equal \u03b8 j : \u2225\u2225R \u03b1x R \u03b2y R \u03b3 z b j + [ x, y, z] T \u2212 a j \u2225\u22252 \u2212 \u03b82 j = 0 (9) where (see Fig. 8 ): b j are the coordinates of the centers B j of the spherical joints S referred to frame , a j are the coordinates of the centers A j of the universal joints U referred to frame W, and R vw denotes a 3 \u00d7 3 rotation matrix of angle v about axis w . If the left-hand side of Eq. (9) is denoted by f j , then particularizing j for { 1 , . . . , 6 } yields the constraint function f = [ f 1 , . . . , f 6 ] T of Eq. (2) , which defines the self-motion manifolds in this example. To densely sample the self-motion manifolds at each task point t = [ y, z] T following Algorithm 2 , we proceed as follows",
            " The existence of mechanical interferences between different links is checked using the SOLID library [40] through Java Native Interface. Eight links L j are considered: the six legs, the base, and the mobile platform. All these eight links are modeled as cylinders, as detailed next. Each leg L j ( j = 1 , . . . , 6 ) is modeled as a cylinder with radius 0.025 and axis A j B j , where A j and B j (i.e., the centers of the universal and spherical joints) are the centers of its top and bottom sections (see Fig. 8 ). The fixed base L 0 is also modeled as a cylinder centered at the origin of frame W, with its axis coincident with the Z axis of frame W, with radius 0.65 and height 0.025. Similarly, the mobile platform L 7 is modeled as a cylinder centered at the origin of frame , with its axis coincident with the Z axis of frame , with radius 0.4 and height 0.025. The geometry described in this paragraph and in the previous one is precisely the geometry that can be observed in Fig. 8 . To apply the method described in Section 3 and obtain the workspace interior barriers and boundaries in a region of the task plane ( y, z ), the following parameters are used. Barriers and boundaries are searched in task box T = [ \u22120 . 41 , 0 . 15] \u00d7 [0 . 41 , 0 . 81] . Each axis of this box is discretized into N t i = 400 equally spaced points. Therefore, box T is approximated by a regular grid G of 1.6 \u00b7 10 5 task nodes in which to compute the 1-dimensional self-motion manifolds. The task of computing, clustering, and matching manifolds in all these nodes is distributed over N p = 8 processors working in parallel (the distribution of the workload is done partitioning the y axis, as indicated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000152_boca-2014-0001-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000152_boca-2014-0001-Figure6-1.png",
        "caption": "Figure 6: Effect of enzyme amount on the ester conversion percentage in a solvent-free system: free lipase (FL, empty bars) and immobilized lipase (IL, filled bars). The error bars indicate the standard deviation (2%).",
        "texts": [
            " According to the low biocatalyst amount, the maximum ester yields were ~ 6% and ~ 10% for free and immobilized lipase, respectively, at an acid:alcohol molar ratio of 1:2. Furthermore, the ester yield increased with increasing acid:alcohol molar ratio up to a threshold value. Unauthenticated Download Date | 1/12/15 9:57 PM The effect of enzyme concentration was studied for the 1:2 ratio using free lipase (0.2 mg, 0.4 mg and 0.8 mg) and the same amount of immobilized lipase (corresponding to around 50 mg, 100 mg and 200 mg of the supported enzyme) in a solvent-free system at 40 \u00b0C. Fig. 6 shows the effect of enzyme concentration on ethyl butyrate synthesis after 48, 72, 96 and 120 h. The ester yield increased with the amount of enzyme in the reaction medium for both the free and immobilized lipase. Using 200 mg of the enzyme-support system (equivalent to 0.8 mg of free lipase), the ester conversion percentage was ~ 23%. These results indicate that use of the immobilized enzyme leads to higher ester yields than the same amount of free lipase. Unauthenticated Download Date | 1/12/15 9:57 PM The use of a lipase in organic media, rather than aqueous, shifts the reaction in favor of product formation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003202_j.ijsolstr.2017.10.008-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003202_j.ijsolstr.2017.10.008-Figure11-1.png",
        "caption": "Fig. 11. Segments of equilibrium path for Dome II under p 1 .",
        "texts": [
            " The origin of the cordinate system is in the support plane below the crown, and the nitial coordinates of the joints are listed in Table 6 . The upper-ring oints, lower-ring joints, and support joints, respectively, lie on horzontal circles of radii 25, 50, and 75, have initial heights 20.1, 18, nd 11.5, and lie close to a spherical surface of radius 150. The numerical solution procedure is similar to that for Dome I, ith k = 1 in Eq. (6) and i = 1, 2, \u2026, 19 in Eqs. (7) . Green\u2013Lagrange train is assumed. For loading at the crown, segments of the equiibrium path are plotted in Fig. 11 for the ranges 0 \u2264 \u03b41 \u2264 1.1 and \u2264 p 1 \u2264 60. Data are presented in Tables 7 and 8 , and shapes are epicted in Figs. 12 and 13 for the xz plane with joints 29, 14, 5, 1, , 8, and 20 ( Fig. 10 ). Fig. 8. Equilibrium shapes at A, C, D, F For Fig. 11 , the displacements of the upper-ring joints are symetric about the z axis. The displacements of the lower-ring joints re wavy, with equal heights for alternating joints around the ring. ables 7 and 8 list h 8 and h 9 , and the heights of the other lowering joints can be determined from h 8 = h 10 = h 12 = h 14 = h 16 = h 18 nd h 9 = h 11 = h 13 = h 15 = h 17 = h 19 . The waviness is due to the folowing: (a) some of the outer joints 20\u201337 support one member nd some support two, and (b) members connected to the sup- orts do not all have the same length (e.g., members 8\u201321, 10\u201322, 0\u201324, and 12\u201325 have length 34.78, whereas members 9\u201321, 9\u2013 2, 11\u201324, and 11\u201325 have length 29.52). .1. Unilateral displacement control For unilateral displacement control, as the crown is pushed ownward, the dome follows segment A-B-C in Fig. 11 . The first nap occurs at C where \u03b41 = 0.0972 and the central joint lies in the ame horizontal plane as the upper-ring joints ( Table 7 ; Fig. 12 ). he central joint jumps downward to the unstrained configuration , G, I, and J for Dome I under p 2 . Table 7 Results for unilateral displacement control of joint 1 in Dome II, using Green\u2013Lagrange strain. Point \u03b41 p 1 h 1 h 2 - h 7 h 8 h 9 \u03b5T (%), bar \u03b5C (%), bar A 0 0 20.1 18 11.5 11.5 0, \u2014 0, \u2014 C 0.0972 0 18.147 18.147 11.521 11.477 0.175, 2\u20133 0.176, 1\u20132 D 0",
            " Point \u03b41 p 1 h 1 h 2 - h 7 h 8 h 9 \u03b5T (%), bar \u03b5C (%), bar A 0 0 20.1 18 11.5 11.5 0, \u2014 0, \u2014 B 0.0391 3.710 19.315 18.059 11.520 11.497 0.102, 2\u20133 0.124, 1\u20132 B \u2032 0.2336 3.710 15.404 17.900 11.500 11.521 0.062, 1\u20132 0.083, 2\u20133 E 0.4307 53.735 11.442 16.150 11.745 11.896 0.477, 1\u20132 0.940, 2\u20133 E \u2032 1.0160 53.735 \u20130.321 3.606 10.928 10.970 0.629, 1\u20132 0.281, 2\u20133 H 1.485 306.25 \u20139.757 \u20133.287 7.251 7.577 1.979 1\u20132 2.838, 8\u20139 4 w w a b j p h t h T 5 t 3 r 0 fi p p a n f 5 m t t i D. With further pushing of joint 1, the segment D-B \u2032 -E-F is followed in Fig. 11 (assuming that the assumptions in the analysis are valid). At F, where p 1 = 0 and the inner and outer rings are almost at the same height ( Table 7 ), the dome snaps to the unstrained equilibrium state G with the central portion inverted. As the indentor pushes further downward, the associated force grows large. Although unrealistic, the analysis was conducted to the next critical point (not shown in Fig. 11 ), which is a turning point denoted I with shape shown in Fig. 12 and data given in Table 7 . .2. Force control Under force control, as p 1 is increased, the first snap occurs hen p 1 = 3.710 at point B in Fig. 11 ( Table 8 ). Joint 1 snaps downard to configuration B \u2032 ( Fig. 13 ). When p 1 reaches 53.735 at E, bifurcation point, the plotted equilibrium path becomes unsta- le. The bifurcating path (not shown in Fig. 11 ), which has unequal oint heights around the upper ring, is also unstable and leaves oint E toward smaller values of \u03b41 . At point E, the dome snaps to point E \u2032 , with the lower ring ardly moving. With further increase of p 1 , assuming the assumpions are valid (which is unrealistic after a while), the dome exibits its final snap when it reaches the bifurcation point H in able 8 and shape in Fig. 13 . . Dome II loaded at a ring joint Finally, Dome II is subjected to downward force p 2 at joint 2 in he upper ring"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003356_j.apm.2018.01.018-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003356_j.apm.2018.01.018-Figure7-1.png",
        "caption": "Fig. 7. Boundary conditions for the inflation analysis.",
        "texts": [
            " The 3D representation of the expanded 2D tyre model is in Fig. 6 . The results obtained from the loading and rolling analyses were used to obtain its the elastic strain energy density at one tyre revolution. The 2D axisymmetric tyre model were used to simulate the temperature distribution in the tyre crosssection. Both the the 2D and 3D numerical models were modelled as implicit non-linear structural finite element models. The 2D numerical model developed for the inflation analysis is shown in Fig. 7 . The tyre model was inflated using a Cavity Pressure Load. The Cavity Pressure Load condition enables the user to define the pressure load that can be applied to the specified cavity as a function of time. To close the cavity, a Cavity Surface Element was placed between the inside end points of the tyre. The Cavity was specified with a reference temperature of 298.15 K, a reference pressure of 100 kPa and density of 1.13006 kg/m 3 , which is the density of Nitrogen at the specified temperature and pressure conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002937_ilt-06-2015-0072-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002937_ilt-06-2015-0072-Figure1-1.png",
        "caption": "Figure 1 Journal bearing configuration",
        "texts": [
            " is lubricant viscosity. Constantinescu (1962) indicated values of the coefficients K and Kz according to: K 12 0.026(Re )0.8265 (3) Kz 12 0.0198(Re )0.741 (4) To introduce the following non-dimensional variables in the modified Reynolds equation: x R , z z L , e c , p pc2 UR , h h c , c Equation (1) in the non-dimensional form can be written as: G h , l K P R L 2 z G h , l Kz P z 1 2 h (5) Where: G (h , l ) h 3 12l 2h 24l 3 tanh h 2l (6) The non-dimensional film thickness h is given by (Crosby and Chetti, 2009) (Figure 1): h 1 cos sin sin cos sin (7) Where: 1 is eccentricity ratio based on major clearance; is attitude angle; is ellipticity ratio. The negative and positive signs in equation (7) are used to determine the film thickness in the upper lobe and the lower lobe, respectively, of the journal bearing. Turbulence on the performance of a two-lobe journal bearing Boualem Chetti Industrial Lubrication and Tribology Volume 68 \u00b7 Number 3 \u00b7 2016 \u00b7 336\u2013340 D ow nl oa de d by U ni ve rs ity o f T ec hn ol og y, S yd ne y A t 0 9: 13 2 4 Ju ly 2 01 6 (P T ) The film pressure is ambient at the bearing ends, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure8-1.png",
        "caption": "Fig. 8. Influence of axial error on the contact region.",
        "texts": [
            " It can be found that: (a) when the tangential error is positive, total contact regions will deflect to the root of pinion and the topland of gear in the conjugation process of teeth from the toe to the heel, respectively; (b) the influence of tangential error gradually reduces to zero in TL direction, but the influence of tangential error gradually increases in TH direction; and (c) the influence level of tangential error on the contact region is less than that of center distance error. (3) Influence of axial error Fig. 8 shows the deflection tendency of contact region in TL and TH directions of pinion and gear when the axial error \u03c0z = 0.03. According to Fig. 8, it can be found that: (a) when the axial error is positive, total contact regions will deflect to the root of pinion and the topland of gear when the tooth pair conjugates from the toe to the heel, respectively; (b) the influence of axial error in TL direction gradually reduces to zero, but the influence of axial error gradually increases in TH direction; and (c) comparing with the influences of center distance error and tangential error on the contact region, the influence of axial error is close to that of tangential error, but is less than that of center distance error"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001585_aim.2018.8452392-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001585_aim.2018.8452392-Figure6-1.png",
        "caption": "Figure 6. Unit part in the previous robot",
        "texts": [
            " As a result, the artificial muscle expands in the radial direction and contracts in the axial direction. The contraction force at this time is used as an actuator. Figure 4. Artificial muscle As a previous research, we developed a peristaltic motion robot named PEW-RO IV. A schematic of PEW-RO IV is shown in Fig.5. Table 2 also shows the specifications of the robot. The unit part will be described first. The unit part is an artificial muscle that is mounted on this robot. It contracts and expands, thus resulting in forward motion (Fig.6). The expansion is propagated backward in order to reproduce peristaltic motion, resulting in forward motion inside the pipe. In order to fix the artificial muscle, a holding ring is used. In addition, the retaining ring is fastened with a nut to fix the artificial muscle. By using this method, the binding force against the expansion of the artificial muscle can be enhanced, and slip-off of the artificial muscle can be prevented. The robot also includes a joint portion. This joint has a structure imitating a universal mechanism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure7-1.png",
        "caption": "Fig. 7. Influence of tangential error on the contact region.",
        "texts": [
            " 6(b,d) presents the boundary of 2D contact region, consisted of five instantaneous contact regions, and the deflection values of actual CL in TL and TH directions. It can be found that:(a) relative to the theoretical state, when the center distance error is positive, total contact regions will deflect to the topland of pinion and the root of gear when the tooth pair conjugates from the toe to the heel, respectively; (b) the influence of center distance error on the contact region will gradually reduce in the conjugation process of teeth, and the shift of bearing contact tends to zero in TL direction. (2) Influence of tangential error Fig. 7 shows the deflection tendency of contact region in TL and TH directions of pinion and gear when the tangential error \u03c0y = 0.03. It can be found that: (a) when the tangential error is positive, total contact regions will deflect to the root of pinion and the topland of gear in the conjugation process of teeth from the toe to the heel, respectively; (b) the influence of tangential error gradually reduces to zero in TL direction, but the influence of tangential error gradually increases in TH direction; and (c) the influence level of tangential error on the contact region is less than that of center distance error"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003761_imece2018-86315-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003761_imece2018-86315-Figure5-1.png",
        "caption": "Figure 5: Cavity, rail and additive material model for thermal analysis (finite volume)",
        "texts": [
            " To evaluate thermal load distribution along the cooling time, transient heat transfer analysis is performed using finite volume method. The physical model considered for the analysis is a three-dimensional cubic cavity containing a 150-mm-long rail/additive solid body at the center. The length, width and height of the cavity is triple the length, width and height of the solid body respectively to ensure that the developed boundary layers on the cavity walls will not interrupt the solid region (Figure 5). The additive material is at an initial temperature of \ud835\udc47\u210e = 1800\u2103 and the rail is initially at the room temperature of \ud835\udc47\ud835\udc50 = 25\u2103. The cavity is filled with air acting as the cooling fluid. Four of the cavity walls, which are normal to \ud835\udc66 and \ud835\udc67 directions, are insulated. Respecting the two cavity walls normal to the \ud835\udc65 direction, the left-side-wall is at the constant inlet velocity of 0.5 m/s (the mean indoor air velocity [27]), and the right-side-wall is set at pressure-outlet boundary condition with fixed atmospheric pressure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002132_j.triboint.2014.09.007-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002132_j.triboint.2014.09.007-Figure6-1.png",
        "caption": "Fig. 6. Radial bearing deformation for the converged high deformation condition.",
        "texts": [
            " The converged results for the high deformation condition F ! ext \u00bc 2; 20\u00f0 \u00de KN are illustrated in Figs. 5,6. As can be seen in Fig. 5, the local bearing deformation in this case yielded to a high perturbation in the hydrodynamic pressure field, resulting in a bimodal pressure distribution typically noticed in big-end bearings simulations. Such high perturbation also increased the area of hydrodynamic load capacity and hence clearly reduced the peak pressure and the area of the cavitation zones. In Fig. 6, it is possible to notice the large structural distortions induced by the hydrodynamic pressures, especially those caused due to the high flexibility of the connecting-rod bearing along the direction in which the present load is applied. Regarding the convergence of the EHL solution, again the Inexact Quasi-Newton Method (IQN-ILS) was the best, converging after 51 iterations (212.53 s). Similarly, the Fixed Point Gauss-Seidel Method (PGMF) showed the worst performance, converging completely only for the (fixed) relaxation factor \u03c9EHL \u00bc 0:15 and after 173 iterations (608"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003060_tmag.2017.2704662-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003060_tmag.2017.2704662-Figure5-1.png",
        "caption": "Fig. 5. Flux density distribution of the CSFCL under (a) under normal condition; (b) positive half cycle of a fault; and (c) negative half cycle of a fault",
        "texts": [],
        "surrounding_texts": [
            "To validate the effectiveness of CSFCL, a 220 V model was designed and established in ANSOFT based on the structure proposed above. Parameters of the model are shown in Table I. The CSFCL is set in a simple three-phase system in the simulation setup. The rated voltage of the three-phase power source is 220 V, and the rated normal current of each phase is 10 A. The three-phase system impedance is ZSi=0.1+j22, and the load impedance is ZLi=19  (i=A, B, C). In normal state, the outer limbs are in deep saturation. When a single-phase-to-ground fault occurs, the right outer limb will be in deeper saturation, and the left outer limb will come out of saturation in the positive half cycle. The outer limbs come out of saturation alternately during a fault. B. Analysis of Direct Excitation Current and Energy Consumption As analyzed, the PM used in the CSFCL can reduce the required DC source. Fig.7 shows the relationship between the required direct excitation current and the amount of used PM. It\u2019s obvious that the needed direct excitation current decreases with the increased amount of PM. The energy loss in DC biasing circuit is one of the largest losses of SFCLs. The energy consumption is estimated assuming that each coil has a nominal resistance of 0.1 and a projected lifetime of 30 years (\u2248263000hours) [8]. It can be seen in Fig. 7 that the use of PMs leads to the reduction of energy consumption. 0018-9464 (c) 2016 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. BU-13 4 As the AC fluxes generated by three-phase AC coils are canceled by each other in the common core in normal state, the CSFCL can further reduce the needed biasing MMF. As shown in Fig.7, the required direct excitation current of the CSFCL is lower than that of a single-phase SFCL which has the same iron core with CSFCL. Hence, the superior property of the CSFCL to a single-phase SFCL with reduced energy consumption is validated. C. Analysis of the Clipping Performance From (4), it\u2019s obvious that the value of Rg influences the excitation impedance Zm of the equivalent four-winding transformer. The variation of Zm of CSFCLs with different air-gaps and PMs is shown in Fig. 8. It\u2019s obvious that the Zm decreases significantly with the increase of air-gap length (lg). According to (5), the impedance of CSFCL is related to the value of Zm. Hence, the fault clipping performance is degraded by the increase of lg. Figure 9 shows the A-phase fault currents limited by CSFCLs when lg is different. It\u2019s apparent that the larger airgap leads to the worse clipping performance, although the amount of ferromagnetic material is reduced. Hence, the parameter of air-gap should be designed properly."
        ]
    },
    {
        "image_filename": "designv11_13_0000691_1350650115569553-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000691_1350650115569553-Figure1-1.png",
        "caption": "Figure 1. Schematic diagram of two-layered porous oil journal bearing.",
        "texts": [
            " The aim of the present study is to solve the governing equations for hydrostatic double-layered porous oil bearings with tangential velocity slip and to investigate its effect on load-carrying capacity, attitude angle, friction variable, and volume flowrate. The effect of various parameters on the bearing performance has been investigated and depicted in the form of graphs. A schematic diagram of an externally pressurized double-layered porous oil journal bearing along with the coordinate system used in the analysis is shown in Figure 1. Entire outer surface of bearing is supplied with pressurized oil at a supply pressure ps. The governing equations of pressure in porous layers for an anisotropic bearing in nondimensional form can be written as follows: for the coarse layer Kx2 @2 p02 @ 2 \u00fe R H 2 @2 p02 @ y2 \u00fe Kz2 D L 2 @2 p02 @ z2 \u00bc 0 \u00f01\u00de For the fine layer Kx1 @2 p01 @ 2 \u00fe R H 2 @2 p01 @ y2 \u00fe Kz1 D L 2 @2 p01 @ z2 \u00bc 0 \u00f02\u00de Modified Reynolds equation in the film region considering velocity slip and anisotropic permeability in dimensionless form is @ @ h 3 1\u00fe x\u00f0 \u00de @ p @ \u00fe D L 2 @ @ z h 3 1\u00fe z\u00f0 \u00de @ p @ z \u00bc @ @ h 1\u00fe ox\u00f0 \u00de \u00fe @ p01 @ y y\u00bc1 \u00f03\u00de The boundary conditions used for solving the above differential equations are as follows for the coarse layer p02 \u00f0 , 0, z\u00de \u00bc 1:0 at 04 42 p02 \u00f0 , y, z\u00de \u00bc p02 \u00f0 \u00fe 2 , y, z\u00de at 04 y4 , 14 z4\u00fe1 at UNSW Library on July 16, 2015pij"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001621_rpj-04-2017-0057-Figure17-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001621_rpj-04-2017-0057-Figure17-1.png",
        "caption": "Figure 17 Machining toolpath for surface finishing",
        "texts": [
            " The following section will provide implementation and testing examples. The methods presented are implemented in a C11 program which is available as an installable toolbar within the MasterCAM software package and has been tested on metal AM components. The part in Figure 16 was manufactured through EBM inTi-6Al-4Vwith amachining allowance of 0.05 inches (1.27 mm), while CNC milling was conducted on a HAASVF2ss. Example toolpaths for planar and cylindrical surfaces within MasterCAM are shown in Figure 17. Recall, the colored D ow nl oa de d by T he U ni ve rs ity o f T ex as a t E l P as o A t 0 4: 24 1 6 O ct ob er 2 01 8 (P T ) feature mesh model is used extensively for analysis and process planning; however, all toolpath generation is done on the native CAD file in the CAM package. Hence, the accuracy of the machining process is not inherently different than conventional NC programming; at least not due tomodel input. A dimensional inspection was conducted for the machined example part using a ZEISS CalypsoTM CMM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002619_s12555-014-0457-x-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002619_s12555-014-0457-x-Figure4-1.png",
        "caption": "Fig. 4. Isothermal region division.",
        "texts": [
            " 3 shows that the heating rate on the leading edge increases along with the angle of attack. Besides the increasing angle of attack and the constantly moving stagnation point, the wing would also compress the incoming flow severely, thus, the heating rate on the leading edge increases continuously. According to the calculations on the above 96 operating conditions, the reentry vehicle can be divided appropriately into several isothermal regions based on the heating rate on the surface of the vehicle (see Fig. 4), and a database of aerodynamic characteristics can be set up to support the following trajectory optimization in Section 3. Robust Adaptive Fault-Tolerant H\u221e Control of Reentry Vehicle Considering Actuator and Sensor Faults Based on ... 201 The space vehicle would experience severe aerodynamic heat and large overload during the reentry flight. So reentry trajectory optimization should be studied so as to improve thermodynamic environment of the reentry vehicle. 3.1. Multiple constraints Equation (1) can be described as X\u0307 = f(X,U, t), (6) where X = [V,\u03b3,\u03c8,x,h,z]T and U = [\u03b1 ,\u03b5]T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003553_s1560354718040081-Figure18-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003553_s1560354718040081-Figure18-1.png",
        "caption": "Fig. 18. Trajectories corresponding to the parameter values (2.10) and \u0393 = 15.03: a) the phase trajectory in the space (p1, p2, M); b) the trajectory of the elliptic profile.",
        "texts": [],
        "surrounding_texts": [
            "2.4. Attracting Tori and the Neimark \u2013 Sacker Bifurcation\nAs noted above, in addition to attracting cycles, the system under consideration can exhibit attracting tori. In this case, one observes a supercritical Neimark \u2013 Sacker bifurcation, which leads to the birth of an invariant curve from a fixed point on the Poincare\u0301 map (2.5). In the chart of dynamical regimes (Fig. 6a), invariant curves correspond also to black regions (i. e., regions without attracting points), which in the chart of Lyapunov exponents (see Fig. 6b) are colored white (that is, the largest Lyapunov exponent is zero). Specifically, in Fig. 6 this is a region on the level \u0393 \u2248 40, and the corresponding invariant curve is presented in Fig. 5b. For a better illustration of the appearance of an invariant curve, we plot a bifurcation diagram on the plane (p2, \u0393) (see Fig. 16) for the following parameter values of the system:\nA = 1.25, B = 5, I\u0303 = 0.705, a = 0.334, b = 1, \u03a9 = 20, \u0393 \u2208 [14.8, 15.5], \u03bc1 = 0.1, \u03bc2 = 0.18, \u03bc3 = 0.705.\n(2.10)\nIt is seen from Fig. 16 that, as the circulation increases above some limiting value \u0393\u2217 \u2248 15.0316, the fixed point of period 2 disappears on the Poincare\u0301 section. Also, an attracting set arises on the section. This set consists of a pair of closed curves (see Fig. 17) to which an invariant torus corresponds in the phase space, that is, a Neimark \u2013 Sacker bifurcation occurs. As the parameter \u0393 increases further, these curves merge to form a closed curve (see Fig. 17). This means that a bifurcation inverse to the torus-doubling bifurcation occurs.\nThe phase trajectories and the trajectory of the profile which correspond to different attracting sets in Fig. 17 are shown in Figs. 18\u201320.\nREGULAR AND CHAOTIC DYNAMICS Vol. 23 No. 4 2018",
            "To complete the picture, we also present the spectrum of Lyapunov exponents for trajectories on the Poincare\u0301 map in a neighborhood of the above-mentioned limiting sets.\nREGULAR AND CHAOTIC DYNAMICS Vol. 23 No. 4 2018",
            "The periodic point with \u0393 = 15.03 corresponds to\n\u03bb1 = (\u22121.27 \u00b1 0.59) \u00b7 10\u22123, \u03bb2 = (\u22121.39 \u00b1 0.66) \u00b7 10\u22123, \u03bb3 = (\u2212294.01 \u00b1 0.25) \u00b7 10\u22123. (2.11)\nThe invariant curve with \u0393 = 15.0339 corresponds to\n\u03bb1 = 0 \u00b1 0.23 \u00b7 10\u22122, \u03bb2 = (\u22122.09 \u00b1 0.21) \u00b7 10\u22122, \u03bb3 = (\u221227.61 \u00b1 0.14) \u00b7 10\u22122. (2.12)\nThe invariant curve with \u0393 = 15.057 corresponds to\n\u03bb1 = 0 \u00b1 0.15 \u00b7 10\u22123, \u03bb2 = (\u221229.40 \u00b1 0.89) \u00b7 10\u22123, \u03bb3 = (\u2212267.6 \u00b1 1.4) \u00b7 10\u22123. (2.13)\n2.5. Strange Attractors\nIn the chart of dynamical regimes presented in Fig. 6a, four extensive regions (black) stand out where there are no limit cycles, and, according to Fig. 6b, the largest Lyapunov exponent in these regions is positive. This is indicative of chaotic behavior of the system at these parameter values.\nREGULAR AND CHAOTIC DYNAMICS Vol. 23 No. 4 2018"
        ]
    },
    {
        "image_filename": "designv11_13_0001686_2.0101816jes-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001686_2.0101816jes-Figure1-1.png",
        "caption": "Figure 1. Schematic of the phosphate electrode. The sharp tip of a tungsten wire forms the sensitive part of the sensor. The wire is mechanically stabilized and insulated in a quartz tube filled with AB glue and a copper wire at its other end is used for measuring the sensor potential.",
        "texts": [
            " Its mathematical expression is: E = E0 + S\u2217lgX Where E is the potential value of the sample solution to be measured, E0 is the standard electromotive force of the electrode; S is the slope; and X is the ion concentration of the solution to be measured. If the ionic strength of the test solution is high enough and constant, the electrode potential E has a linear relationship to the logarithm of the ion concentration in solution. Tungsten rod preparation.\u2014Pure tungsten rod (99.99%, 1 mm in diameter) was cut into 4 cm long segments and the metal surface was burnished by coated abrasive (800#, 1600#, 2500#) to form a sharp tip at one end (Fig. 1). The segments were cleaned with deionized water ) unless CC License in place (see abstract).\u00a0 ecsdl.org/site/terms_use address. Redistribution subject to ECS terms of use (see 132.174.254.155Downloaded on 2018-11-29 to IP using the ultrasonic cleaning device (Kunshan Ultrasonic Instruments Co., Ltd) and air-dried. Assembling the prepared tungsten rod.\u2014The prepared tungsten rod was inserted into a quartz tube (1.5 mm in diameter, Guanghua Quartz Factory Co., Ltd), revealing the electrode tip (about 15 mm length) and tail (about 5 mm length). The electrode tail was welded to a copper wire and the junction was wrapped in insulating tape. The gap between the electrode and the quartz tube was filled with high strength AB glue (T300, Dongguan Weimao Industrial Investment Co., Ltd) to prevent the penetration of sample solution and to fix the tungsten rod in the quartz tube (Fig. 1). Potential measurement.\u2014No pretreatment of the electrode was required before measurements. The potential difference between the working electrode (tungsten) and the reference electrode was measured using the CS2350 controlled bipotentiostat (Wuhan Corrtest Instruments Co., Ltd, China) at a constant temperature of 26 \u00b1 1\u25e6C. The pH was adjusted with 10\u22122 M Sodium hydroxide or 10\u22122 M Hydrochloric acid.35 The sensor performance, including selectivity, response time, reproducibility, detection limit and life time were examined by measuring the voltage potential for a HPO4 2\u2212 concentration range of 10\u22127-10\u22121 M at a pH of 10 and were accomplished in accordance with the recommendations of the International Union of Pure and Applied Chemistry (IUPAC)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001462_aac3cb-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001462_aac3cb-Figure10-1.png",
        "caption": "Figure 10. Forces at the highest points. The arrows in cyan indicate the additional acceleration due to the rotation of the circle as the pendulum starts moving down.",
        "texts": [
            " This angular acceleration influences the whole circle of riders, and size of the effect depends on the position in the circle. For riders in the highest position of the circle this leads to an additional acceleration in the z (\u2018vertical\u2019) direction with the Phys . Educ . 53 (2018) 045017 8July 2018 size |r\u03b8\u0308| = | \u2212 rg sin \u03b8/L|. In the turning points, where |\u03b8| = \u03b80 = 120\u25e6, this gives a contribution (r/L)g sin \u03b80 \u2248 g \u221a 3/12 \u2248 0.14g. For a rider in the lowest point of the circle, the angular acceleration of the pendulum gives an equally large, but negative, acceleration in the z direction. Figure\u00a010 shows these additional accelerations that depend on the location in the circle and account for the differences in the normal forces on the rider, seen in figure\u00a06. For high school students, understanding the forces for a non-rotating pendulum may be sufficiently challenging. The data collected by the four students in different positions of the rotating pendulum illustrate that the ride experience depends on the position, which can be quite confusing. This paper is intented to provide teachers with background to discuss the small deviations when student questions arise"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002720_978-3-319-24055-8-Figure5.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002720_978-3-319-24055-8-Figure5.1-1.png",
        "caption": "Fig. 5.1 Whining noise generation process",
        "texts": [
            " STE\u00f0h1\u00de \u00bc Rb2:h2\u00f0h1\u00de Rb1h1 \u00f05:1\u00de where Rbj, is the base radius of gear j. It is mainly due to voluntary (corrections) and involuntary (defects) geometrical deviations of the teeth at a micrometric scale and to elastic deformation of loaded teeth, wheel bodies and crankshafts. STE fluctuations also generates mesh stiffness fluctuations. Under operating conditions, the parametric excitations induce dynamic loads at the gear meshes, which are transmitted to the gearbox receiving structure via the wheel bodies, crankshafts and bearings, as presented in Fig. 5.1. The vibratory state of the crankcase is the main source of the radiated noise [3]. 64 A. Carbonelli et al. STE fluctuations need to be minimized by introducing voluntary tooth micro-geometrical modifications in order to reduce the radiated noise. For this study, the selected optimization parameters for each gear pair are: \u2022 the tip relief values X of pinion and driven gear i.e. the amount of material remove on the teeth tip, \u2022 the starting tip relief diameters / of pinion and driven gear, \u2022 the added up crowning centered on the active tooth width Cb;i=j"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000477_195102-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000477_195102-Figure1-1.png",
        "caption": "Figure 1. Schematic diagram describing the fabrication of the control channels on graphene and the motility control experiment using an electric field on it (not to scale). (a) Transfer of graphene onto a glass and its functionalization with 1-pyrenebutanol (Py\u2013OH). (b) Patterning SU-8 polymer on graphene by means of photolithography. A ring-shaped graphene surface with a 15 \u00b5m width was exposed. (c) Coating the substrate with kinesin motor protein. (d) In vitro motility assay on the graphene-based fluidic channel. A positive bias voltage was applied to a graphene electrode to control the motility of microtubules.",
        "texts": [
            " Microtubule motility was observed with a fluorescence microscope (Nikon, TE2000U) equipped with a 100\u00d7/1.40NA oil objective, an EMCCD (Nikon, DQC FS) and a LED light source (Custom Interconnect Ltd, CoolLED). A CoolLED pE-1 excitation system (Custom Interconnect Ltd, CoolLED) with an excitation wavelength of 550 nm and a TRITC filter (Nikon, EX 540/25, DM 565, BA 605/55) was used as a light source to illuminate rhodamine-labeled microtubules. We analyzed the trajectories of the microtubules using the Metamorph analysis software (Molecular Devices). Figure 1 displays our processes to fabricate a control channel on graphene and perform motility assay on it. Detailed procedures are described in section 2. Briefly, multilayer graphene was prepared and transferred onto a glass substrate as reported previously [26]. Then, we placed the graphene substrate in 1 mM solution of 1-pyrenebutanol (Py\u2013OH) to functionalize the graphene surface with negatively charged hydroxyl groups (figure 1(a)) [28, 29]. A metal electrode (Ti/Au) was fabricated on the edge of the graphene via thermal evaporation (figure 1(b)). A SU-8 photoresist layer with ring-shaped clearing was patterned on the functionalized graphene using photolithography [30]. This ring-shaped artificial structure was chosen to enable easy distinction of the region of interest. Since the photoresist pattern worked as an insulating layer, the electric field from the graphene can be localized on the ring-shaped exposed graphene region. A flow cell composed of the graphene electrode, 0.1 mm thick double-sided tapes and a counter-electrode was constructed. For in vitro motility assays, the flow cell was incubated with a casein solution so that the casein layer adsorbed onto the graphene worked as a passivation layer. Then, it was subsequently incubated with a kinesin solution, resulting in a kinesin-adsorbed graphene surface (figure 1(c)). The motility solution containing rhodamine-labeled microtubules (tubulin concentration: 0.1 \u00b5M) and 5 mM ATP was added and incubated [31]. The two electrodes were connected to a DC power supply (E3631A, from Agilent Inc.) and positive bias voltages (in the range of 0\u20133 V versus the counter-electrode) were applied to the graphene electrode to drag negatively charged microtubules on the exposed graphene region (figure 1(d)). The motions of the microtubules were observed via fluorescence microscope (Nikon TE2000U) through the transparent graphene electrode. Here, it should be noted that the functionalized graphene was utilized as a transparent and conducting substrate for both observation and electrical control of the motion of the microtubules. To construct a graphene-based biomotor system, graphene should support the motility of a kinesin\u2013microtubule system. Figures 2(a) and (b) show the fluorescence micrographs of microtubules on a bare graphene substrate and a Py\u2013OH functionalized-graphene substrate coated with kinesin, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002737_9781118899076-Figure5.2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002737_9781118899076-Figure5.2-1.png",
        "caption": "FIGURE 5.2 The three-electrode cell configuration typically used for protein film electrochemistry experiments with a PGE working electrode surface. The electrode\u2013 potentiostat interface is denoted by asterisks.",
        "texts": [
            " Instead, we focus here on the wealth of biochemical information to be gained purely from the electrochemical experiment and illustrate this with examples drawn from our own research interests. However, first, we outline some of the practicalities to be considered when performing PFE of enzymes. PFE is typically performed using a cell configuration incorporating reference, counter, and working electrodes. One example of a glass electrochemical cell frequently used for PFE in our laboratories is illustrated in Figure 5.2. The reference electrode is housed in a side arm connected to the sample chamber by a Luggin capillary that minimizes physical mixing of the solutions in the sample and reference chambers while defining a sensing point for the reference electrode near the surface of the 107THE FILM ELECTROCHEMISTRY EXPERIMENT working electrode. The sample chamber contains the platinum wire counter electrode and is shaped to minimize solution turbulence during rapid working electrode rotation as is frequently required during studies of redox enzyme catalysis",
            " The working electrode is prepared from a cylinder (3 mm diameter) of the desired electrode material mounted onto a brass rod with silver-loaded epoxy and subsequently sealed into a nylon sheath with epoxy resin. The sheath allows the working electrode to be mounted onto the shaft of a rotor that makes electrical contact with the working electrode via the brass rod. During experiments, the working electrode is positioned in the sample chamber at the level of the Luggin capillary and with its exposed face in contact with a 3.5 ml solution (Fig. 5.2). The entire electrochemical apparatus is placed inside a Faraday cage within a N2-filled chamber to minimize electrical noise and the presence of oxygen, respectively. The temperature of the sample chamber is maintained by means of a jacket filled with circulating water under thermostat control. Several electrode surfaces have been found to support nondestructive yet robust protein adsorption [1\u20133]. We favor either the edge or basal plane of pyrolytic graphite. These surfaces have been successful in adsorbing a variety of enzymes such that they retain similar redox and/or catalytic properties to those described in solution, which are the usual criteria for confirming the functional integrity of adsorbed protein",
            " The optimum composition of the solution for enzyme adsorption is usually found by exploring a range of ionic strengths, pH, buffer, and electrolyte identity 108 ENZYME FILM ELECTROCHEMISTRY in addition to enzyme concentration and where the inclusion of a coadsorbate, for example, positively charged neomycin, may help to secure an electroactive protein film. The time invested in exploring these conditions is rewarded when it comes to performing and analyzing experiments that probe enzyme function since both are greatly facilitated by the reproducible formation of stable films. Once the film is prepared, its catalytic performance can be readily assessed in a range of conditions using the cell illustrated in Figure 5.2. The design allows aliquots of reagents to be injected directly into the experimental solution from concentrated stock solutions, or the sample solution can be removed completely and replaced with a fresh solution of choice. Modification of the basic cell design allows for rapid equilibration of the sample chamber with gaseous agents of varying partial pressures [10]. In essence, the enzyme film can be subject to instantaneous dialysis by a variety of means and to great advantage as we illustrate later"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003504_taes.2018.2852419-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003504_taes.2018.2852419-Figure2-1.png",
        "caption": "Fig. 2. Engagement geometry",
        "texts": [
            " Thus we have x\u0307T = vT cos\u03c8T (9) y\u0307T = vT sin\u03c8T (10) \u03c8\u0307T = \u03c9, (11) 0018-9251 (c) 2018 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. where \u03c9 := aT /vT . In the practical application, if \u03c9 is unknown to M, a state estimator is required to estimate \u03c9 online, the feasibility of which is tested in Section VI. A schematic view of planar interception geometry is shown in Fig. 2. The objective of the guidance law design is to intercept T with a desired impact angle \u03b3c, which is referred to as a perfect interception. Mathematically, we need to guarantee that after tgo, three relations are required: xM \u2212 xT = 0, yM \u2212 yT = 0 and \u03b8 \u2212 \u03c8T = \u03b3c. For the practical use of the guidance law, we also require the maneuver effort to be zeroed after a finite time, because the stored fuel is quite limited for an exoatmospheric missile. The zero-effort impact angle, which is the impact angle obtained by setting u \u2261 0, is denoted by \u03b3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002445_ijmmme.2016010103-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002445_ijmmme.2016010103-Figure5-1.png",
        "caption": "Figure 5. Gearbox surface temperature variation- Isothermal gear oil bath temperature (800C)",
        "texts": [
            " In this research work efforts have been made to measure the surface temperature variation of gearbox surface due to gear oil temperature, frictional condition and dynamic loading under the influence of SAE 80W-90 lubricating gear oil. The properties of SAE 80W-90 are- density 887 kg/ m3 (15.60C), viscosity 139cSt (400C), 15 cSt (1000C), Viscosity Index 110, flash point 2180C, pour point (-270C). Finite element analysis used for the structural and thermal simulation of transmission gearbox. The internal temperature of gearbox has an influence on thermal stresses and deformations leads to failure. Figure 5 shows the temperature variation at different point of gearbox surface for gear oil bath temperature of 800C. Figure 5(a) highlights the temperature variation in gear train assembly at the value of h100w/m 2 k for convective heat transfer co-efficient (h). Temperature variation is shown in colour code. The minimum temperature is 342.94 k and maximum is 350.01 k. Maximum temperature effect is found on counter shaft right end in red hues, which was fixed in simulation. Red hues signify the maximum temperature change and thermal stress generation portion. In between minimum and maximum temperature 8 another temperatures are shown in figure. The temperature varies very gradually at different points, it can be seen from figure 7. In figure 7, h100 (1) shows the gradual increment in temperature at different points of gear train. Around second gear pair the temperature and deformation is high. Figure 5(b) shows the temperature variation in gear train assembly at the value of h200w/m 2 k. As the value of h increases there is very small change in temperature profile of gear train. The minimum temperature decreases 0.09 k and increase in maximum temperature of 1.58 K. The temperature profile varies between (342.85-351.59) k. The high temperature region is at same place of fixed portion of right side end of counter shaft. Figure 5(c) shows the temperature increment on 3rd gear. The temperature reached to 350.01 k. On 2nd gear the temperature is 349k under the influence of h300w/m 2 k. The temperature profile varies between (342.93-352.03) k. Figure 5(d) shows the gear train temperature profile varies (342.92-352.38) k. First gear is in loose meshing and its temperature is 347.12 k and 2nd, 3rd gear shows temperature increment and have a value of (349.23-351.33) k. The lay/counter shaft also shows variation of temperature (346.07-352.38) k for h400w/m 2 k. Figure 5 (e, f) shows very similar temperature profile for gear train. The 3rd gear profile shows temperature increment in red hues and this area is prone for thermal stresses more for (h500w/m 2 k, h600w/m 2 k) convective heat transfer coefficient values. In figure 7, h400 and h500 shows the gear train temperature profile variation. For second part of study the gear oil temperature is increased to 1000 C. Figure 6 shows the temperature profile of gearbox surface at gear oil temperature of 1000C. Figure 6(a) shows the temperature profile of gear train assembly at h100w/m 2 k convective heat transfer coefficient (h) value",
            " The right side fixed counter shaft end temperature is maximum for h300w/m 2 k. The temperature profile varies between (342.8-370.09) k. In figure 8, h300 (3) shows the temperature linear temperature variation at different point on gear train. Figure 6 (d, e, f) shows gradual change in temperature. Gear profile of 2nd and 3rd shows red hues of hot areas. In figure 8, h600 shows the gear train temperature profile variation. Multi speed transmission gearbox and gear oil analysis is highly nonlinear problem. From FEA analysis (Figure 5 and 7), when h is 100 w/m2k and gear oil temperature is constant at 800C, the gear train profile temperature varies very gradually and linearly. As we increase the value of convective heat transfer coefficient to 200 W/m2 k the difference in maximum temperature is only 1.58 k. Further as we increase the value of h (300,400,500 & 600) the difference in gearbox surface temperature profile varies only (1.58-2.7) k (Figure 7). Same temperature profile as figure 7 is generated when gear oil temperature is constant at 1000C and value of h increases from (100-600) W/m2 k, the difference in temperature is (1-8) k (figure 8). From figure 7 and figure 8 it can be concluded that if gear oil temperature is constant (isothermal) the gear train thermal stresses at each convective heat transfer coefficient value is within permissible limits. The temperature profile of gear train shows that the lower temperature is approximate constant (Figure 5 & 6) and maximum temperature is varying by (1-5) k (Figure 7 & 8). Convective heat transfer co-efficient (h) refer to transfer of heat with fluid movement. In general for the increase value of h the rate of heat transfer increases. It signifies that the gearbox surface temperature should be reduced with increase in h value. The numerical simulation results show the same results as per thermal concept. In this research work for thermal analysis gear oil temperature was varied for 800C and 1000C and h varied for (100-600) W/m2k, for these loading conditions the temperature at 10 different gear assembly points was measured",
            " The present research work of multi speed transmission gearbox has analytical and theoretical significance for performance improvement of gearbox. The FEA simulation results of medium duty transmission gearbox concludes that the gradual increment in temperature around engaged gear pairs reported the effect of heat generation (thermal stresses) on surface of gears in influence of gear meshing, frictional heat, average heat transfer and gear oil bath temperature. Thermal stresses are within permissible limits (Figure 5 & 6). The research work suggest that if the overheating of gear oil is less, which refer cooling will be more and it increase the thermo-mechanical performance of multi speed transmission system. The simulation result shows the equivalent von-mises stresses and thermal strain in permissible range for the transmission gearbox. These simulation results help in preliminary design stage to predict the thermal nature of transmission gearbox and for the design optimization of transmission parts. FEA numerical simulation results are in agreement with experimental results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001085_tmag.2017.2708140-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001085_tmag.2017.2708140-Figure5-1.png",
        "caption": "Fig. 5. Distribution of eddy current density by using 3D FE method under ac load condition considering tapped holes.",
        "texts": [
            " (16), the eddy current loss occurring in the bolt fixing the magnet can be considered. D. Analysis result based on ac load condition Fig. 4 shows the phase current waveform produced by the PMLSG obtained under ac load condition. As shown in Fig. 4(a), the effective value of the current is at rated velocity 1.6 m/s. As shown in Fig. 4(b), the Fast Fourier Transform (FFT) analysis results can be obtained from the phase current. In the case of the PMLSG with rated ac load, the phase current contains 1st, 3rd, 5th and 7th order harmonics, and the dominant harmonic is 1st order. Fig. 5 and 6 show the distribution of eddy current density obtained from the 3D FE method and 3D analytical method, for the case when the phase current waveform is the same as Fig. 4. The end effects were verified from results of the FE and analytical methods. Fig. 7 show the eddy current density according to each harmonic order for the PMLSG with phase current waveform of Fig. 4, and also show the eddy current density according to analysis regions. The results of the 3D analytical method agreed with the FE method results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002541_j.jsv.2016.01.019-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002541_j.jsv.2016.01.019-Figure7-1.png",
        "caption": "Fig. 7. First modes of the elastomer sample: (a) Mode 1, f\u00bc1152 Hz; (b) Mode 2, f\u00bc1320 Hz; (c) Mode 3, f\u00bc1359 Hz; (d) Mode 4, f\u00bc1551 Hz.",
        "texts": [
            " In particular the Young modulus value is taken from measurements at a frequency (2 kHz) close to the expected eigenfrequencies. Constraints with null displacement are imposed to model the attachment at the sample holder. Free boundary conditions are imposed elsewhere including the sliding contact line. The absence of stick phase in measured signals leads to a friction force which is only velocity-dependent. The contribution of an elastic component of the friction force is negligible in this case. The first four eigenmodes and their natural frequencies are shown in Fig. 7. Mode 1 corresponds to a longitudinal deformation of the base along the contact line (x-axis). This deformation being normal to the sliding velocity, this mode cannot be excited. Mode 2 is a deformation of the sample along the y-axis. As the sliding velocity is also oriented in the yaxis, this mode can of course be excited by the friction force. Mode 3 corresponds to a twist deformation around the vertical z-axis. A friction force oriented in the y-direction and uniform along the contact line cannot excite this mode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure12-1.png",
        "caption": "Fig. 12. Undercutting \u2014 type IIb: z=6; x=xmin=0.449; \u03b4r\u204e=0.068; \u03b4t\u204e=0.114; \u03b1=20\u00b0; ha\u204e=0.8; c*=0.7; \u03c1*=10.",
        "texts": [
            " Besides, the starting point b of the involute profile ba lies on the base circle, and as a result, the teeth are not undercut in a radial direction (\u03b4r=0 mm, \u03bbr=0%). In this case only a tangential undercutting is obtained, caused by the rack-cutter fillet AF, where \u03b4t=0.47 mm and \u03bbt=4.82%. This is determined by the fact that the rack-cutter fillet (profiled over a circle of a radius \u03c1*>\u03c1max\u204e=2.05) is placed in the area ADE (Fig. 7), outside the boundary area ACE. When generating the teeth, shown on Fig. 12, the parameters of the gear and rack-cutter, excluding the coefficient \u03c1*=10, are the same as on Fig. 11. And in this case, from Fig. 12 it is seen that the teeth are not undercut of type I, as the tip line g\u2212g passes through point A'. The undercutting obtained is of type IIb and is also provoked by the rack-cutter fillet AF. In this case, due to the larger value of \u03c1*, the fillet curve AF is positioned outside both boundary areas ACE and ADE. As a result, the teeth are undercut in a tangential, as well as in a radial direction and the undercutting indices are respectively: \u03b4r=0.68 mm; \u03bbr=4.76%; \u03b4t=1.14 mm; \u03bbt=12.21%. Tooth undercutting is also observed on metal prototypes of gears, shown on Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001853_s10846-015-0233-z-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001853_s10846-015-0233-z-Figure9-1.png",
        "caption": "Fig. 9 The sensor housing, a CAD model design and b fabricated",
        "texts": [
            " The Gumstix tiny computer acts as the high-level controller of the robot and communicates with the SSC-32 using UART via the Robostix expansion board. The Gumstix runs Linux 2.6 and is connected to the Netpro-VX expansion board for wireless connectivity. The housing for the Gumstix and expansion boards were fabricated using a rapid prototype machine. To physically measure the normal foot forces, the Lynxmotion hexapod robot, shown in Fig. 7, was equipped with Force-Sensitive Resistors (FSR-402) similar to [27]. The sensors were calibrated after they were embedded into the rapid prototyped housings, as shown in Fig. 9. The sensor data is read by the Robostix, shown in Fig. 7, and used for real-time monitoring of the stability. 4.2 Stability Metric A stability criterion needs to be defined within the controller to give the robot a sense of stability, to measure the stability of the robot, and to predict instability. Theoretically, any of the existing stability margins may be used for this purpose. However, the need for calculating the stability of the robot at any instant requires high frequency calculations within the controller which indicates a need for a concise stability margin with low computation cost and low sensor input information"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000739_s0263574713000829-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000739_s0263574713000829-Figure1-1.png",
        "caption": "Fig. 1. (Colour online) The Sarcos Primus humanoid robot (a) and its simplified model (b).",
        "texts": [
            "cambridge.org Downloaded: 28 Jun 2014 IP address: 155.198.30.43 where Imax is the impulse of the biggest push that the robot can stand and M is the total mass of the robot. For normalization, the robustness criterion is found as Krobust = Imax \u2217 TC M \u2217 Lff , (4) where Tc = \u221a HCoM/g. For the push recovery methods based on LIPM, their maximum robustness value is 1. Using (4), we compare the robustness of four typical methods, which is shown in Table I. The Sarcos Primus humanoid robot, as shown in Fig. 1(a), is a hydraulic force-controlled humanoid robot. When building its model in the simulator (see Fig. 1b), its upper body is simplified to focus on its integrated mass\u2013inertia properties. The Young\u2019s modulus\u2013coefficient of restitution element is used instead of the spring\u2013damper model as the contact model between the feet and the ground in the simulator.14 There are seven joints in each leg: two in the hip, one in the thigh, one in the knee, one in the shank and two in the ankle. The thigh joint is fixed, so qrl \u2208 R6\u00d71 and qll \u2208 R6\u00d71 represent the angles of the other six joints of the right leg and left leg, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002668_1350650115619610-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002668_1350650115619610-Figure1-1.png",
        "caption": "Figure 1. (a) Schematic diagram of elastohydrostatic thrust bearing system having elliptical shape recess; (b) flexible thrust pad hydrostatic bearing.",
        "texts": [
            " Four different types of geometric shapes of recess have been used in the present study to show the influence. The fluidfilm pressure distribution, pocket pressure, fluid-film stiffness, and damping coefficient have been compared for a wide range of tilt parameter and restrictor design parameter. The result of this paper will help the bearing designer and academic community to understand the effect of elastic deformation on bearing performance. Analysis A schematic diagram of a deformable hydrostatic thrust pad bearing with elliptical shape recess is shown in Figure 1. The present analysis involves interaction of deformable structure with an internal lubricating flow. This type of problem cannot be solved without the use of iterative method. Therefore, Reynolds equation coupled with elasticity equation has been solved by using a nonlinear iterative finite element method used in Figure 2. Lubricant flow field equation The governing Reynolds equation used to compute fluid-film distribution and other performance parameter is expressed as follows.28,35\u201340 Governing flow field equations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001043_s11044-017-9572-9-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001043_s11044-017-9572-9-Figure2-1.png",
        "caption": "Fig. 2 Illustration of approximation of planar contact as multiple point contacts: (top) foot is only partially in contact with the environment; (bottom) the complete foot is in contact with the ground",
        "texts": [
            " Constraint on the contact wrench acting on the robotic foot, which needs to be fulfilled in order to ensure stability of the contact, will be derived using the span representation (5) of the friction pyramid. Sliding conditions are included, and this model also considers the \u201cyaw friction\u201d (the torque generated by friction in the direction of a normal z-axis [4]), which is crucial for avoiding undesirable yaw rotations of the robot\u2019s feet. In general, surface contact will be any contact between an arbitrary curved ground and planar foot where at least three non-collinear point contacts exist (Fig. 2). The surface contact of a single foot and the ground is modeled by multiple separate point contacts placed at each corner of the foot (for a rectangular foot, four contact points placed at the corners). It is important to note that the proposed procedure for deriving contact wrench conditions can be applied when a number of contact points is greater than one. When there are two contact points, we have line contact. If we have 3 or more non-collinear contact points, then there is surface contact. The convex hull of the contact points defines the contact perimeter. To show that the proposed procedure does not require any shape of the foot, we will model it as a pentagon of arbitrary shape (Fig. 2). Total contact wrench acting on the robot foot, calculated for the reference point P (inside support area), is [ FP MP ] = [ \u2211n i=1 Fi\u2211n i=1 ri \u00d7 Fi ] = [ \u2211n i=1 U\u03b1i\u2211n i=1[ri]\u00d7U\u03b1i ] . (7) Vectors r1 to rn represent positions of contact points relative to reference point P . [u]\u00d7 is a skew-symmetric cross-product operator matrix. Matrix U can be also written in the form U = \u23a1 \u23a3 \u03bc cos \u03c0 m . . . \u03bc cos \u03c0(2m\u22121) m \u03bc sin \u03c0 m . . . \u03bc sin \u03c0(2m\u22121) m 1 . . . 1 \u23a4 \u23a6 = \u23a1 \u23a3 \u03bcC \u03bcS 1 \u23a4 \u23a6 , (8) where 1 represents row vector of length m with all elements 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002962_mfi.2016.7849485-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002962_mfi.2016.7849485-Figure1-1.png",
        "caption": "Figure 1. Basic structure of a quadrotor",
        "texts": [
            " Compared with the neural network discussed above, our scheme uses random optimization of weights, which is easy to implement in real time within the UAV and shows good performance according to the simulation results. The rest of this paper is organized as follows. Section II describes the basic dynamic model of the quadrotor. Section III presents the architecture of the NN augmented dynamic inversion model as applied to a quadrotor. Section IV provides result and discussions. Section V concludes this work and outlines future work. A. The Quadrotor UAV A quadrotor is a multirotor helicopter that produces lift with four fixed-pitched rotors in a plane. Fig. 1 shows a basic model of an unmanned quadrotor. It generally uses two pairs of identical propellers: the front and rear motors rotate counter clockwise (CCW) while the other two rotate clockwise (CW). By changing the speed of each rotor it is possible to specifically generate a desired total thrust. Lifting and landing movement is achieved by increasing or decreasing all the four motors. Roll movement is obtained by varying the speed of the left and right motors. Pitch movement is obtained similarly using the front and rear motors",
            " And yaw movement is controlled by increasing (decreasing) the speed of the front and rear motors and decreasing (increasing) the speed of the lateral motors. The quadrotor is under-actuated since it has six degree of freedom (three rotational and three translational) but only four actuators. Therefore, only four of the six DOF can be controlled and stabilized. B. Flight Dynamics Consider an earth-fixed inertial frame E={Oexeyeze}and a body-fixed frame B={Obxbybzb}whose origin O is at the center 978-1-4673-9708-7/16/$31.00 \u00a92016 IEEE 174 of mass of the quadrotor, as shown in Fig. 1.Defined p = (x, y, z)T to represent the position of the quadrotor with respect to the inertial frame and , , T to represent the attitude of the quadrotor. Furthermore, V = (u, v, w)T describes the linear velocity of the quadrotor in the body frame and =(p, q, r)T representing the rotational velocity of the quadrotor within the body frame. According to Newton s law, the kinematics and dynamics equations can be described as ( ) ( ) b b b bi bi bi b b b bi bi bi dP F mV mV dt dL M I I dt (1) where b bi is a skew-symmetric matrix and defined as follows: 0 0 0 b bi r q r p q p (2) Separately, for translational motion, the combined force can be denoted by (in body fixed frame) |= - -b b b i T D b i GF F F C F (3) where 2 2 2 2 1 2 3 4 0 0 0 , , 0 dx b b i T D dy G i dzT k x F F k y F k z mgc Thus the translational motion equation can be described as 1b b b b b b bi T D bi bi u V F F g V v m w (4) | b i b bi x y C V z (5) where m denotes the quadrotor mass, and g is the gravity acceleration, and | cos cos cos sin sin cos sin sin sin cos cos cos sin sin sin sin cos sin sin cos sin cos sin cos cos sin sin cos cos b iC For angular motion, the moment can be denoted by (in body fixed frame) = - T D b b bM M M (6) Define the thrust coefficient and the drag coefficient as Tc and Dc , if the distance from the rotors to the centre of mass is denoted by d, then the control torques generated by the four rotors are 2 2 2 4 2 2 3 1 2 2 2 2 2 4 1 3 T T D dc dc c (7) Considering the gyroscopic effects and disturbances,then the torques are 2 2 2 4 1 2 3 4 2 2 1 3 1 2 3 4 2 2 2 2 1 2 3 4 30 30T T T m b T T m Q Q Q Q dc dc J Q M dc dc J Q c c c c (8) D rx b ry rz k M k k (9) Thus the angular motion equation can be described as 1 D b b b b b bi T bi bi p I M M I q r (10) 1 ( ) ( ) ( ) ( ) 0 ( ) ( ) 0 ( ) / ( ) ( ) / ( ) b bi t s t c p c s q H s c c c r (11) Define the state vector X=[p q r u v w x y z]T,the input vector U=[ 1 2 3 4]T, and the system modeling equations can be finally derived as 1 | , 1 D b b b b T bi bi b bi b b b b b T D bi bi b i b bi p q r I M M I H f X U X u F F g V m v C V w x y z (12) A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000929_b978-0-444-52215-3.00008-8-Figure8.2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000929_b978-0-444-52215-3.00008-8-Figure8.2-1.png",
        "caption": "FIGURE 8.2 Schematic diagram of a heat exchanger. A cold fluid flows in the inner tube or tubeside and a hot fluid flows in the outer tube or shell-side. Heat is transferred by (i) convection from the hot flowing fluid to the outer surface of the inner tube, (ii) conduction from the outer tube surface to the inner tube surface, and (iii) convection from the inner tube surface to the flowing cold fluid. The hot fluid can be placed on the tube-side and the cold fluid can be placed on the shell-side. The choice of tube-side or shell-side for a given fluid depends on design objectives as discussed in the text and in the caption of Figure 8.3. Image adapted from www.wikipedia.org.",
        "texts": [
            " One can also analyze the heat losses of the hot block. Sometimes the energy transport is transient or time-dependent such as in the example of Figure 8.1 in which a hot block is placed on a cold block. Other times, energy transport can be steady-state in which temperature gradients between two bodies are maintained by constant supply and constant removal of heat. The focus of this chapter is on steady-state energy transport. In this chapter, the device introduced for transferring energy from one fluid to another fluid is the heat exchanger (see Figure 8.2). The main heat transfer modes of a heat exchanger are forced convection, due to movement of two fluids exchanging heat, and conduction, due to physical barriers that are used in the heat exchanger to separate the two fluids at different temperatures. Introduction to Supercritical Fluids560 A heat exchanger is the most common type of device that is used to transfer heat between two flowing fluid streams (Figure 8.2). Heat transfer occurs by convection via fluid flow through the device and conduction via the heat exchanger tube wall. In Figure 8.2, a cold fluid at temperature Tc1 enters the heat exchanger on the tube-side and exits the heat exchanger at temperature Tc2 on the tube-side. A hot fluid enters the heat exchanger on the shell-side and exits the heat exchanger on the shell-side. The temperature of the cold fluid increases and the temperature of the hot fluid decreases. The cold fluid receives heat, \u00feq, from the hot fluid, which causes the temperature of the cold fluid to increase from Tc1 to Tc2 and the hot fluid loses exactly the same amount of heat, q, which causes its temperature to decrease from Th1 to Th2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002746_978-3-319-44156-6_8-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002746_978-3-319-44156-6_8-Figure3-1.png",
        "caption": "Fig. 3 Transition configurations of the variable-DOF 7R mechanism: a Trans. conf. I; b Trans. conf. II; c Trans. conf. III; d Trans. conf. IV; e Trans. conf. V; f Trans. conf. VI",
        "texts": [
            " Substitution of this pose into the remaining hyperplane equations Hi(v1, v3), Hi(v1, v4), Hi(v6) and Hi(v7) yields linear equations in the remaining unknown tangent half of the joint angles. This algorithm yields the input-output equations of all joint parameters vi, i = 2,\u2026 , 7 against the input parameter v1 in the arbitrary 7R motion. In case of \ud835\udf0b-turns of the joints, one should check when the leading coefficients of the equations E1, E2 and P1 vanish identically with respect to each unknowns separately. This yields the following four special configurations v1 = \u221e, v2 = 0, v3 = \u221e, v4 = \u221e, v5 = 0, v6 = \u221e, v7 = 0 (Fig. 3a), v1 = \u221e, v2 = 0, v3 = \u221e, v4 = 0, v5 = \u221e, v6 = 0, v7 = \u221e (Fig. 3b), v1 = 0, v2 = \u221e, v3 = 0, v4 = 0, v5 = 0, v6 = \u221e, v7 = 0 (Fig. 3c), v1 = 0, v2 = \u221e, v3 = 0, v4 = \u221e, v5 = \u221e, v6 = 0, v7 = \u221e (Fig. 3d). (13) As it will be shown later, through these configurations, the 7R mechanism can switch from the 1-DOF 7R modes to the 2-DOF double-Bennett mode. In this section, we will reveal that P2(v1, v2) = 0 in Eq. (12) corresponds to a 2-DOF double-Bennett mode (Fig. 1a). Solving the linear equation P2(v1, v2) = 0 with respect to v2, and substitution into \ud835\udc29 we obtain \ud835\udc29b = ( \u2212( \u221a 2 \u2212 1)v5, \u2212( \u221a 2 \u2212 1)v5, 1, \u22121, \u2212( \u221a 2 \u2212 2)v5, ( \u221a 2 \u2212 2)v5, \u221a 2, \u221a 2 )T . (14) It is noted that this point is independent of the input joint angle v1",
            " (12) have common solutions, which are geometrically the intersections of the curves in Fig. 2. These common solutions can be found using an algebraic approach as follows. Substituting the solution of P2(v1, v2) = 0 with respect to v2 into P1 = 0, one obtains a univariate polynomial equation in v1. It has the two conjugate complex double solutions \u00b1I and the two real solutions vT11 = \u22123 7 \u221a 7 \u2212 1 7 \u221a 7 \u221a 2, vT21 = 3 7 \u221a 7 + 1 7 \u221a 7 \u221a 2. (16) The corresponding values for v2 and v5 to the real solutions of v1 are vT12 = \u22122 7 \u221a 7 \u221a 2 \u2212 1 7 \u221a 7, vT15 = 3 7 \u221a 7 + 1 7 \u221a 7 \u221a 2 (Fig. 3e), vT22 = 2 7 \u221a 7 \u221a 2 + 1 7 \u221a 7, vT25 = \u22123 7 \u221a 7 \u2212 1 7 \u221a 7 \u221a 2 (Fig. 3f ). (17) Using Eq. (15) the remaining joint angles can be computed. Equations (16) and (17) show that in these two transition configurations (Fig. 3e, f), v5 = \u2212v1 holds. As one can see in Fig. 2, the two discrete 1-DOF 7R modes are connected via the 2-DOF double-Bennett mode through the six transition configurations (Fig. 3). This paper has presented a new variable-DOF 7R mechanism. The algebraic analysis of the 7R mechanism has shown that the 7R mechanism has two 1-DOF 7R modes and one 2-DOF double-Bennett mode and it can switch from the two 1-DOF 7R modes to the 2-DOF double-Bennett mode via transition configurations. In the future, the construction approach in this paper and [5] will be used to gen- erate more new variable-DOF mechanisms using four-link mechanisms. 1. Bennett, G.: A new mechanism. Engineering 76, 777\u2013778 (1903) 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure24-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure24-1.png",
        "caption": "Fig. 24. Typical Stress distribution due to tip corner contact.",
        "texts": [
            " The variation of the stress along the major axis of the contact pattern is uneven in the annulus-planet tooth pairs, and more uniform in the sun-planet, as the similar results show in Figs. 17 and 18 . Because of the curvature relation of concave-convex contact, the minor axis length of the annulus-planet tooth pairs is larger than that of the sun-planet, but the contact stress is smaller. The contact stresses are analyzed for various contact positions before the beginning and after the end of normal contact. A typical 3D stress distribution of those analysis results is shown in Fig. 24 . Fig. 25 shows the variation of the stress distribution along the minor axis at the various contact positions for both the gear pairs. The tooth contact with concentrated stresses occurs on the tip corner of one engaged gear flank, while the contact position varies within the root area of another gear flank. Two different conditions are involved in the calculation by using the proposed LTCA approach: without and with consideration of tip corner contact beyond the normal line of action. The influences of tip corner contact on the shared loads and the contact stresses are discussed as follows"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure15-1.png",
        "caption": "Fig. 15. The boundary condition of the planetary gear.",
        "texts": [
            " 14 (c) and (d). The influences of the twist of the sun gear shaft and the deformation of the carrier are involved in the FEM analysis. The element type is Hex8 (6 faces and 8 nodes). No special elements are necessary to be placed at the points of contact for solving the contact problem by using the FEM software (MSC.Marc), because the special algorithm, so-called Solver Constrain Method, is applied. The FE-model includes 577,970 elements and 715,730 nodes. The setting of the boundary conditions is shown in Fig. 15 . The transmitted torque is applied on the face-end of the sun gear shaft. Each planet gear is connected to the axis of the planet shaft with rigid link elements, so as to rotate freely around the planet shaft. The nodes either on the output end of the carrier shaft or on the outer circumference of the annulus gear are all defined as fixed. The Young\u2019s modulus and Poisson\u2019s ratio of all the components is equal to 206 GPa and 0.3, respectively. In order to compare with FEM, the distributed contact stress along the face-width and the load sharing among planets are calculated by the proposed LTCA approach and FEM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003448_j.procir.2018.03.277-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003448_j.procir.2018.03.277-Figure1-1.png",
        "caption": "Fig. 1 Example of photorealistic picture of assembled vise.",
        "texts": [
            " Presentation of case study In the course \u00abComputer aided design\u00bb, the students learn to draw 3D-parts, assemble the parts into a model with natural movements, animate movements of the parts and create photorealistic pictures. They also learn to create 2Dproduction drawings with measurements, tolerances etc. During the course, the students have a compulsory assignment which is divided in two subtasks: In task 1 they 3D-model the parts of a vise based on 2D-drawings which are not optimally made. The parts are assembled to a vise with natural movements, and the students create some photorealistic pictures (Fig. 1). The vise is also parameterized, so that the length and the width can easily be changed. In subtask 2, the students also create 2D-production drawings with dimensions, tolerances etc. The drawings are supposed to be more optimally made than the original drawings. An example of one of the original drawings is shown in Fig. 2. The red text on the drawing are comments on aspects the students need to pay attention to and/or improve. This could be missing or not optimally placed dimensions, Thomas Haavi et al",
            " Presentation of case study In the course \u00abComputer aided design\u00bb, the students learn to draw 3D-parts, assemble the parts into a model with natural movements, animate movements of the parts and create photorealistic pictures. They also learn to create 2Dproduction drawings with measurements, tolerances etc. During the course, the students have a compulsory assignment which is divided in two subtasks: In task 1 they 3D-model the parts of a vise based on 2D-drawings which are not optimally made. The parts are assembled to a vise with natural movements, and the students create some photorealistic pictures (Fig. 1). The vise is also parameterized, so that the length and the width can easily be changed. In subtask 2, the students also create 2D-production drawings with dimensions, tolerances etc. The drawings are supposed to be more optimally made than the original drawings. An example of one of the original drawings is shown in Fig. 2. The red text on the drawing are comments on aspects the students need to pay attention to and/or improve. This could be missing or not optimally placed dimensions, Author name / Procedia CIRP 00 (2017) 000\u2013000 3 comments on tolerances and recommendations on how to place dimensions when the parts are parameterized"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003156_s10798-017-9420-5-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003156_s10798-017-9420-5-Figure1-1.png",
        "caption": "Fig. 1 Framework of engineering design curriculum (Fan and Yu 2016)",
        "texts": [
            " 11) also pointed out that the design cycle of a high school engineering project consists of nine steps: identify the need or problem; research the need or problem; develop possible solutions; select the best possible solution; construct a prototype; test and evaluate the solution; communicate the solution; redesign; and finalize the design. S.-C. Fan et al. Framework of high school level engineering design curriculum In our previous study (Fan and Yu 2016), we developed a framework for high school level engineering design curriculum and to guide the development of teaching and learning activities in this study (as shown in Fig. 1). According to this framework, the ring gear guides the completely engineering design process, while the center gear provides the learning context. Meanwhile, scientific inquiry, mathematical analysis, and technology techniques are used as the connection gears between design ideas and implementation. McCormick (2004) noted that the engineering design project learning might fail if teachers do not provide good explanations and learning experiences between conceptual knowledge and problem-solving processes",
            " There were three ten-grade classes; with a total sample of 103 students aged between 16 and 17 years old participated. The selected school recruits students who attain the top 10\u201315% scores on the Basic Competence Test for Junior High School Students. It would be an advantage for implementing engineering design curriculum but also a limitation of this study. Care had been taken to explain the data and findings. An engineering design module, based on the framework of the engineering design curriculum (see Fig. 1), was developed focusing on integrative STEM knowledge and engineering design. The design of this module also referred to the \u2018\u2018Activity-, Project-, and Problem-Based Model (APB Model) (Grimm 2010)\u2019\u2019. As Kolmos et al. (2009) highlighted that a well-designed PBL curriculum should aligned with seven elements, which include: methodological objectives and interdisciplinary knowledge, diverse types of open-end problems and projects, clear progression and course size, constructed students\u2019 learning, well training academic staff and facilitation, supporting space and organization and, formative assessment and evaluation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003448_j.procir.2018.03.277-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003448_j.procir.2018.03.277-Figure3-1.png",
        "caption": "Fig. 3 Photograph of vise no. 1. Made by mechanical engineer students.",
        "texts": [
            " Practical results of the process The assignment was not carried out exactly as planned. The students were supposed to get the nylon parts before they started, but because the nylon printer had a failure and needed to be repaired, the students got the parts at the end of the project period. These are the white parts in Fig. 4 and Fig. 5. Nevertheless, all the groups managed to complete the assignment within the deadline, and ended up with a fully functional vise. The vices made by the mechanical engineer students are shown in Fig. 3 and Fig. 5. The vise made by the design students is shown in Fig. 4. The design students used this unforeseen opportunity to experiment and then print the missing parts in plastic with the cheap desktop printers, as shown in Fig. 6. They also managed to print the small screws in plastic, although steel screws were handed out (see the two screws at the bottom in Fig. 6). Beforehand it was considered unlikely that the threaded parts could be printed in plastic, but the design students did it anyway",
            " Practical results of the process The assignment was not carried out exactly as planned. The students were supposed to get the nylon parts before they started, but because the nylon printer had a failure and needed to be repaired, the students got the parts at the end of the project period. These are the white parts in Fig. 4 and Fig. 5. Nevertheless, all the groups managed to complete the assignment within the deadline, and ended up with a fully functional vise. The vices made by the mechanical engineer students are shown in Fig. 3 and Fig. 5. The vise made by the design students is shown in Fig. 4. The design students used this unforeseen opportunity to experiment and then print the missing parts in plastic with the cheap desktop printers, as shown in Fig. 6. They also managed to print the small screws in plastic, although steel screws were handed out (see the two screws at the bottom in Fig. 6). Beforehand it was considered unlikely that the threaded parts could be printed in plastic, but the design students did it anyway",
            " They made drawings of ten parts plus a drawing of the assembly. An example of one of the drawings is shown in Fig. 7. It is not possible to know how the quality of the drawings would have been if the students had not taken the optional additional work with 3D-printing. However, it should be noted that the 14 students participating had from none (0!) up to max three comments on their set of 2Ddrawings, which is an indication of an increased level of understanding when comparing with those students not taking on this task. Fig. 3 Photograph of vise no. 1. Made by mechanical engineer students. Fig. 4 Photograph of vise no. 2. Made by design students. Fig. 5 Photograph of vise no. 3. Made by mechanical engineer students. Fig. 6 The additional parts printed by the design students. Author name / Procedia CIRP 00 (2017) 000\u2013000 5 3.2. Cognitive result; students\u2019 learning outcome Results are presented in accordance with the structure of the interview guide presented in 2.2. The project as a whole: The two engineering groups both said that the time allocation was more than sufficient, and had it not been for printing time, this part of the assignment could have been completed in one day"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002441_iros.2015.7354054-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002441_iros.2015.7354054-Figure5-1.png",
        "caption": "Fig. 5. A 3-dimensional example is shown where the quadrotor is steered towards a goal point through a window on a slanted wall. The window has tight clearance with respect to the robot. The height of the window is only 25% larger than the robot\u2019s diameter, however, the diameter is a conservative estimate provided by the minimum-radius bounding sphere. The width of the window is 75% larger than the robot\u2019s diameter.",
        "texts": [
            " As can be seen, the algorithm can have large deviations from the deterministic results in the presence of uncertainty while still avoiding collisions. The maximum deviations were observed to correlate with the robot taking a more conservative trajectory around the ends of the obstacles and the deviation grew over time as the deterministic case results in a faster completion of this trajectory. A simulation was performed where the quadcopter was guided through a window-like opening in a large wall (see Fig. 5). The goal position of the robot was set directly on the other side of the window from the quadcopter\u2019s initial position. The quadcopter strafes along the non-vertical wall and then passes through the window when it reaches it and is not avoiding collisions with the wall in front of it. This simulation demonstrates the algorithms capabilities to perform 3- d collision avoidance with non-vertical obstacles even when using a simple 1-d approximation of the uncertainty. Videos of the above experiments along with other scenarios can be found at the University of Utah DARC Lab webpage1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001357_acc.2018.8431879-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001357_acc.2018.8431879-Figure1-1.png",
        "caption": "Fig. 1. UAV configuration",
        "texts": [
            " The article [7] develops an estimation algorithm for the wind velocity using the inversion of the equations that define the components of the accelerometer measurements. The paper outline is as follows. In Section II the considered UAV drone is described and the flight dynamics model is derived. A preliminary study of the dynamics is carried out in Section III. The wind estimation algorithms are presented in Section IV. The results of numeric experiments are shown in Section V. The remarks and discussion conclude the paper in Section VI. This section presents the model of the UAV dynamics, which has configuration as it is shown in Fig. 1. According to the identification work at low/medium velocity in [9] using the Parrot AR Drone 2.0 and similar, rotors 978-1-5386-5428-6/$31.00 \u00a92018 AACC 1921 gyroscopic effects and inertial counter torques are neglected since they are rather small. The translational dynamics of the drone in the body frame yield mu\u0307+m\u03d6 \u00d7u = Faero +mRT g, (1) where m is the mass of the UAV, u = [u v w]T is its linear velocity expressed in body frame, \u03d6 = [p q r]T is its angular velocity in body frame, Faero = [FXaero FYaero FZaero] T is the vector of the external aerodynamic forces in body frame, g = [0 0 g]T is the gravity acceleration in inertial frame, R is the rotational matrix defining the passage between the inertial and body frames"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure3-1.png",
        "caption": "Fig. 3. Schematic illustration of directrix, meshing line and datum surface.",
        "texts": [
            " Referring to the rack-cutter profiles of W\u2013N gear in Chinese standard, the tooth profile of HNCGs with point contact is standardized as Table 1. 2.3. Derivation of datum surface and directrix In HNCGs, the datum surface is a hyperboloid of one sheet, on which the directrix of tooth surface lies, and the set of conjugated points of directrixes of pinion-gear tooth surfaces in space is a straight line, which is called as the meshing line in the nomenclature of this novel gear. Thus, both the directrix of tooth surface and its meshing line can be regarded as the generatrix generating the datum surface, as show in Fig. 3. The meshing line Cm is represented in S1 by the vector function Cm : \u03c1 \u00bc \u03c1 u\u00f0 \u00de \u00bc R1i1 \u00fe u sin \u03b4 j1\u2212 cos \u03b4k1\u00f0 \u00de : \u00f04\u00de By the transformation of coordinate system S1 to S2, Cm can be represented in S2 by the vector function Cm : \u03c1 \u00bc R2i2 \u00fe u cos \u03b4 j2\u2212 sin \u03b4k2\u00f0 \u00de : \u00f05\u00de According to the generation rule of datum surface of HNCG, it can be generated by the rotation of the meshing line or the directrix of tooth surface in the three-dimensional space about the axis of pinion and gear, respectively. So, datum surfaces \u03a3p, \u03a3g can be represented in S1, S2 by vector functions \u03a3p : Pp \u00bc B \u2212\u03bb1\u00f0 \u00de\u03c1 u\u00f0 \u00de \u00bc R1e \u2212\u03bb1\u00f0 \u00de \u00fe u sin \u03b4e1 \u2212\u03bb1\u00f0 \u00de\u2212 cos \u03b4k1\u00f0 \u00de 0 \u2264 \u03bb1 b 2\u03c0\u00f0 \u00de \u00f06\u00de \u03a3g : Pg \u00bc B \u2212\u03bb2\u00f0 \u00de\u03c1 u\u00f0 \u00de \u00bc R2e \u2212\u03bb2\u00f0 \u00de \u00fe u cos \u03b4e1 \u2212\u03bb2\u00f0 \u00de\u2212 sin \u03b4k2\u00f0 \u00de 0 \u2264 \u03bb2 b 2\u03c0\u00f0 \u00de \u00f07\u00de When the parameters u and \u03bbi (i = 1, 2) satisfy the relationship u = \u03bb1R1 sin \u03b4 = \u03bb2R2 cos \u03b4, Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001084_j.mechmachtheory.2017.05.017-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001084_j.mechmachtheory.2017.05.017-Figure13-1.png",
        "caption": "Fig. 13. The influence of an intentional tooth modification and loads on an actual contact ratio.",
        "texts": [],
        "surrounding_texts": [
            "To get more insights into the characteristics of tooth modifications based on compensated conjugation presented in this paper, the wave forms and variations of LTE with multi loads are shown in Figs. 14 and 15 . In comparison, the case of gear pairs with a second order parabolic TE is given, which represents the case of traditional convex modification of tooth surface and is widely used [4,30] . Generally, the contact ratio of a gear pair will be reduced after the tooth modification. Hence, before discussing numerical examples, some definitions of actual contact ratio \u03b5 r and geometrical contact ratio \u03b5 t are illustrated as follows: { \u03b5 r = \u03d5 r / \u03d5 p \u03b5 t = \u03d5 t / \u03d5 p (14) where, \u03d5r is the rotational angle that a tooth pair is actually in contact with each other due to the tooth modification and loads; \u03d5p is the rotational angle of the driving gear; \u03d5t is the rotational angle from P 1 to P n . The \u03b5 t is determined by gear dimensions. The \u03b5 r is determined by the values of the tooth modification and the applied loads, and it increases with the applied load due to the tooth deformation. Thus, the actual contact ratio \u03b5 r could be distinguished from the geometrical total contact ratio \u03b5 t . Fig. 14 shows an example of calculated LTE for Gear pair No. 1. The LTE changed its wave form according to the instantaneous number of meshing tooth pairs. Fig. 14 (a) illustrates the case of LTE variations without any modification. At the top, the red curves are the TE curves without any load, which are illustrated three times by shifting the angular pitch. It is can be observed that the curve always consists of two regions of two pairs and three pairs alternately with applied loads. The actual contact ratio \u03b5 r of this case equals to geometrical contact ratio \u03b5 t . Because the tooth deformation is smaller at the region of three pairs than at the region of two pairs and has no intentional TE compensating, the PPTE increased with the load. Fig. 14 (b) shows the case of variations of LTE with the ideal TE. Different to the expected decrease of the tooth contact ratio after the tooth modification, the LTE curves still consisted of two regions of two pairs and three pairs alternately with applied loads, which indicates that the actual contact ratio \u03b5 r of this case is still equal to the geometrical contact ratio \u03b5 t, and the contact ratio doesn\u2019t decrease after the tooth modification. Then, an outstanding advantage of the modification with the ideal TE can be found herein. As shown in Fig. 14 (b), the loads are applied from 0 to 3500 Nm; when loads are applied from 0 to 1200 Nm, the LTE at the region of 2 pairs is less than that of 3 pairs, and the PPTE decrease with the load; at the load of 1200 Nm (the designed input torque), the PPTE is reduce to almost zero, which means that the uneven conjugation due to the tooth deformation is well compensated by the introduced ideal TE, and the exact correction needed for a perfect transmission is achieved; after 1200 Nm, the LTE at the region of 2 pairs is greater than that of 3 pairs, and the PPTE increase with the load, which is resulted by the possibility that the tooth deformation becomes so significant with the load increase that the modification values with the ideal TE are not enough for the compensated conjugation. Furthermore, when the load changed from 900 to 1500 Nm, PPTE is no more than 1 \u2032\u2032 , indicating that the reduction in PPTE is significant over a wide range of loads, and such a characteristic could be another advantage of this tooth modification method. Fig. 14 (c) illustrates the case of variations of LTE with the further longitudinal modification. After a further longitudinal modification, the region of one pair meshing appears. When the load is applied form 0 to 300 Nm, the LTE is through the regions of 1 and 2 pair meshing which means that the actual contact ratio \u03b5 r is greater than 1 and less than 2 before the load of 300 Nm. Meanwhile, the reduction in PPTE is observed to be insignificant from 0 to 300 Nm, and this is because that the introduced TE is specially designed for the case of contact ratios greater than two. When the load increases from 300 Nm, the actual contact ratio \u03b5 r is greater than 2. The overall trend of the variations of PPTE with the load changes little, and the minimum PPTE is still at the load of 1200 Nm. After a further longitudinal modification, the LTE is still effectively compensated by the intentional TE. Fig. 14 (d) shows the case of variations of LTE with a second order parabolic TE, which is used to illustrate the influence of the traditional convex modification on the actual contact ratio \u03b5 r , and the influence of \u03b5 r on the compensated conjugation. Reduction in PPTE at lower loads (almost before 600 Nm) is significant because the actual contact ratio is less than 2. As can be see clearly in Fig. 14 (d) that before the load of 600 Nm, one gear tooth need to go through a \u201ctwo pairs - one pair - two pairs\u201d meshing process, and the tooth deformation is smaller in the two-pair region than in the one-pair region, therefore, the conjugation difference can effectively be compensated by the traditional convex TE. With the load increases, the actual contact ratio \u03b5 r increases, gradually approaching to the geometrical contact ratio \u03b5 t , then, the corresponding wavelike LTE is difficult to be compensated by the TE of traditional convex modification. Fig. 14 (e) shows the comparisons of PPTE under multi-loads between four cases. The case with the ideal TE has the smallest PPTE at the load of 1200 Nm. For the case after a further longitudinal modification, the PPTE slightly increases at 1200 Nm, but the overall trend changes little. For the case without modification, the PPTE increases almost linearly with the increasing load. For the case with the second-order parabolic TE, the reduction in PPTE at lower loads (almost before 600 Nm) is significant. Fig. 15 shows the variations of multi-load TEs for Gear pair No. 2. The PPTE is smaller of Gear pair No.2 than that of Gear pair No.1 because the wave form of the LTE is flattened due to the increase in the number of the meshing tooth pairs. From the results of Figs. 14 and 15 , it is expected that the gear tooth modification method based on compensated conjugation presented in this paper is robust, not limited to the geometrical contact ratio of gear pairs; even the geometrical contact ratio become higher, an excellent result of a lower PPTE can be still obtained."
        ]
    },
    {
        "image_filename": "designv11_13_0000599_j.engfailanal.2012.02.008-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000599_j.engfailanal.2012.02.008-Figure8-1.png",
        "caption": "Fig. 8. The field of equivalent stresses on pedestal traverse.",
        "texts": [
            " Subsequently, it was found out that one bolt was destroyed and in bodies of several bolts were identified cracks. These failures were repaired \u2013 cracks were abraded and welded (Fig. 6), damaged bolts were replaced by new ones (bolts M48 were replaced by bolts M52). At the same time was made decision that it is necessary to accomplish stress analysis of traverse of casting pedestal. Computational model of traverse (due to symmetry only one half of beam and middle part with arm of toothing was considered) is given in Fig. 7. Computation of traverse was accomplished for all load cases given above. In Fig. 8 is given field of equivalent stresses on traverse of casting pedestal for the most danger state of loading. In Figs. 9 and 10 are shown details of equivalent stresses on upper side of traverse beam. For experimental determination of time-dependent changes of stress in the traverse beams was used the electrical resistance strain-gage method [3,6,8]. Strain-gage measurements were realized by strain-gage apparatus SPIDER with application of strain-gages produced by company HBM according to Fig. 11"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002536_0954406215618424-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002536_0954406215618424-Figure4-1.png",
        "caption": "Figure 4. External gear tooth geometry for estimation of deformations.",
        "texts": [
            " Normally, in case of rigid rimmed gears (rim thickness more than five times of tooth height), the last two deformations do not affect the total deformations as much as the first two deformations. However, the last two deformations would not be neglected in the present investigation, as the variation of rim thickness is still considered for analysis. With the above mentioned considerations, the total deformations of a tooth under a normal load F, can be written as T \u00bc B \u00fe S \u00fe G \u00fe C \u00f03\u00de Deformations of external gear tooth. Referring to the external gear tooth geometry shown in Figure 4, various components of tooth deformations of engaged tooth of external gear under the action of normal load F can be expressed as1 The bending deflection BE \u00bc 12F cos2 h EBt3b Y2 i \u00fe h2 3 Yih \u00fe 6F cos2 \u00f0w h\u00de3 EBt3b w Yi w h 4 w Yi w h 2 loge w Yi w h 3 \u00f04\u00de where w \u00bc Htb hta tb ta and h \u00bc rb rd \u00f05\u00de The shearing deflection, SE \u00bc 2\u00f01\u00fe \u00deF cos2 EBtb h\u00fe \u00f0w h\u00de loge w h w Yi \u00f06\u00de at UNIV CALIFORNIA SAN DIEGO on March 12, 2016pic.sagepub.comDownloaded from The deformation along the line of action due to distortion of rim under load i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000776_s11771-012-1135-x-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000776_s11771-012-1135-x-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of bottom following: Reference frames",
        "texts": [
            " Neglecting the equations in sway, roll and yaw, the simplified equations for surge, heave and pitch can be written as [14\u221215] 11 prop 1 22 2 33 prop 3 wq u uq w q m u m wq d X m w m uq d m q d M                    (3) where 11 22 33 2 2 2 2 | | | | | | | | | | 2 | | , , , | | u w y q wq wq uq uq u qq uu ww w q q w q w w ww q q q uq w q uw ww w wu w m m X m m Z m I M m m X m m Z d X q X u X w d Z q q Z w q Z w w Z w d M q q M uq M w q M uw M u w M w M w w                                               The solution to the problem of bottom following proposed here is built on the following intuitive explanation [16\u221217] (see Fig. 1): A simple bottomfollowing controller should compute: 1) the distance between the vehicle\u2019s center of mass Q and the closest point P on the Path 1, and 2) the angle between the vehicle\u2019s total velocity vector \u03bdt and the tangent to the Path 1 at P, and both reduce to zero. This motivates the development of the \u201ckinematic\u201d model of the vehicle in terms of a Serret-Frenet frame F that moves along the path; F plays the role of the body axis of a \u201cvirtual target vehicle\u201d that should be tracked by the \u201creal vehicle\u201d",
            " In this work, however, a Frenet frame F that moves along the bottom path to be followed is used with a significant difference: the Frenet frame is not attached to the point on the path that is closest to the vehicle. Instead, the origin OF=P of F along the path is made to evolve according to a conveniently defined control law, effectively yielding an extra controller design parameter. As will be seen, this seemingly simple procedure is instrumental in lifting the stringent initial condition constraints that are presented in Ref. [18] for path following of marine vehicles. J. Cent. South Univ. (2012) 19: 1240\u22121248 1242 Consider Fig. 1, where P is an arbitrary point on the Path 1 to be followed. Associated with P, consider the corresponding Serret-Frenet frame F. The signed curvilinear abscissa of P along the path is denoted as \u03bc. Clearly, Q can either be expressed as q=[x, 0, z]T in U or as [xF, 0, zF]T in F. Stated equivalently, Q can be given in (x, z) or (xF, zF) coordinates. Let cos 0 sin 0 1 0 sin 0 cos F F F F              R be the rotation matrix from U to F, parameterized locally by the angle \u03b8F. Define F Fq   , then c c c ( ) ( ) ( ) F Fq c c g                (4) where cc(\u03bc) and gc(\u03bc) denote the path curvature and its derivative, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003283_0954406217745336-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003283_0954406217745336-Figure1-1.png",
        "caption": "Figure 1. The structure of bearing\u2013wheel\u2013rail system.",
        "texts": [
            " Then, based on the coupling dynamics theoretical calculation model, the wheel/rail system\u2019s coupled nonlinear dynamic characteristics were studied under random load, and the accuracy and effectiveness of the model were confirmed. Finally, this system was applied to a practical engineering design project. This paper has important theoretical value and practical significance for developing reliable railway bearings and wheel/rail system with good static/dynamic characteristics that can withstand dynamic impact load. The bearing\u2013wheel\u2013rail system consisted mainly of an adapter, bearing, wheel, rail axle, and rail, as shown in Figure 1. The adapter is directly fitted to the bearing, the bearing inner ring and axle are interference fitted, and the wheels are directly fitted onto the corresponding position of the axle, so no slippage was allowed during the operation. The wheel is in direct contact with the rail. The force distribution of the system was calculated according to its structure and working principle and is shown in Figure 2. The adapter was installed between the railway bearings and bogie. It bore the bogie\u2019s impact load and passed that impact load on to the bearing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003019_j.triboint.2017.04.018-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003019_j.triboint.2017.04.018-Figure2-1.png",
        "caption": "Fig. 2. Schematic view of the rolling bearing assembly.",
        "texts": [
            " \u2022 Different additive package, including Zinc, low amounts of Phosphorus and Sulfur and high amount of Calcium instead of Magnesium. \u2022 75W90-A and 75W90-B oils have almost the same physical properties. \u2022 75W85-B oil has the lower viscosity at 40 \u00b0C and 100 \u00b0C and high VI. 3. Experimental procedure for tribofilms generation The total friction torque, generated by the rolling bearing lubricated with each axle gear oil, was measured on a modified Four-Ball machine (Cameron-Plint TE 82/7752), where the four-ball arrangement was replaced by a rolling bearing assembly, developed by Cousseau et al. [20], as shown in Fig. 2. Additional details of this assembly can be found in [20]. This assembly allows to perform tests with several rolling bearing geometries and also to obtain a reliable friction torque and operating temperature measurements. The rolling bearing assembly is composed mainly of two parts. The first part contains the bearing lower race (3) which is fitted in the spacer (2) and this set is fitted on the bearing house (1). These parts of the group (A) are tight clamped to ensure that there is no relative motion between them. The second part (B) consists of the bearing upper race (5) which is mounted on the shaft adapter (6). In operation, the load (P) is applied on the lower plate (13) and the rotational speed (n) is transmitted to the shaft adapter (6), which is connected to the drive shaft of the machine (see Fig. 2). The upper race movement is transmitted to the rollers and cage assembly (4). Then, through the bearing housing (1) the internal bearing friction torque is transmitted to the torque cell (11). The friction torque was measured with a piezoelectric torque cell KISTLER 9339A, ensuring high accuracy measurements ( \u00b1 1 Nmm) even when the friction torque generated in the bearings was very small compared to the measurement range available (see Table 2). During the test, the temperature at several different locations was recorded. Five thermocouples (I\u2013V) were positioned in strategic locations in order to measure and control the lubricant and bearing housing temperatures (see Fig. 2). Two other thermocouples were used to measure the temperatures of the room and of the air flow around the bearing housing. When assembled in the modified four-ball machine, the rolling bearing assembly was exposed to forced air convection to evacuate the heat generated during bearing operation using two 38 mm diameter fans running at 2000 rpm, cooling the chamber surrounding the bearing housing. In order to control the temperature during the tests, the bearing assembly was mounted with two heaters which were controlled with a PID control system with feedback given by thermocouple III (see Fig",
            " After each test, the used oil was collected from the axial rolling bearing test rig and analyzed by Direct Reading Ferrography (DRIII) and Analytical Ferrography (FMIII) techniques. The surface roughness of the raceways was measured and an X-ray photoelectron spectroscopy (XPS) was performed to chemically characterize the tribofilm formed in the rollers of the tested rolling bearing. The surface topography measurements were performed on the lower (stationary) raceway of the roller bearing (see Fig. 2). A 3D optical microscope (BRUKERTM NPFLEX), equipped with a noncontacting optical vertical scanning interferometry (VSI) technique, was used. According to ISO 4288-1996 [28], a filter cut-off value of \u03bb = 0.25 mmc was selected, according to the texture of the rolling bearing raceway. The roughness analysis was performed using a Gaussian filter defined by ISO 11562-1992. The 2D roughness measurements, extracted from the 3D topographies, were measured across the raceway contact track, in the radial direction",
            " An analyzer pass energy of 80 eV is normally adequate for detail scans as it permits to determine the exact position of the peaks. A higher resolution is required and lower pass energies can be used with corresponding loss of signal intensity [30]. The data collected was compared with the peaks curve fitting recorded on the surface of roller bearings raceways submerged in oil without any mechanical work. The Cylindrical Roller Thrust Bearing tests were performed under a constant oil sump temperature of 110 \u00b0C, measured by thermocouple III (see Fig. 2) and used to control the test temperature. Table 4 shows the temperature measured in each test as well as the physical properties of the oil (kinematic viscosity, dynamic viscosity, thermoviscosity, piezoviscosity and Lubricant Parameter) at the operating temperature, as well as the corresponding modified Stribeck parameter Sp and the roller/raceway lubricant center film thickness. The Lubricant Parameter is defined by the product of the dynamic viscosity by the piezoviscosity of the lubricant, at the operating temperature, that is LP \u03b7\u03b1= "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003907_tsmc.2019.2913410-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003907_tsmc.2019.2913410-Figure6-1.png",
        "caption": "Fig. 6. Control signal u.",
        "texts": [],
        "surrounding_texts": [
            "A2 = [\n0.0379 1 \u22120.2846 \u22121\n]\nare introduced, such that the non-\nlinear system (67) can be represented by the following T\u2013S fuzzy system:\nx\u0307 = 2\u2211\ni=1\n\u03b8i(Ai + M Ni)x + B(u + f (x, t)) (73)\nwhere \u03b81 = \u03bcF1 1 (x) = [(4 \u2212 x2 1)/4], \u03b82 = \u03bcF2 1 (x) = (x2 1/4), M = [I I], = diag{\u03bcF1 2 (x)I, \u03bcF2 2 (x)I}, N1 = [ A11 \u2212 A1 A12 \u2212 A1 ] , and N2 = [\nA21 \u2212 A2 A22 \u2212 A2\n]\n. Selecting the parameter of switching\nfunction as G = [0 0.5] and solving the LMIs in Corollary 1, the SMC gains can be obtained\nK1 = [\u22120.9177 \u2212 0.7020], K2 = [\u22120.9493 \u2212 0.8139].\n(74)\nIn order to illustrate the effectiveness of the controller designed method in this paper, the SMC law parameters are\ngiven as \u03b3 = 0.002, \u03b21 = 0.8, and \u03b22 = 0.4 \u221a\n2. For the TSFSS (73), there contain immeasurable premise variables, it is obviously that the existing SMC approach (such as [25], [28], [30], [36], and [39]) which assumed that the premise variables are measurable cannot be used in this system. However, some new matrices are developed to represent the TSFSS by a new ones which the premise variables are totally measurable, and the measurable premise variables in the original system are fully used in the new system. By analyzing the stability of the new system, Theorems 1 and 2, and Corollaries 1 and 2 are obtained, the controller (39) can be designed to stabilize this system. Then, Figs. 2 and 3 show the state and input trajectories, respectively. The sliding mode surface s(t) = 0 is plotted in Fig. 4.\nFrom Figs. 2\u20134, it is shown that the T\u2013S fuzzy system with case 1 is asymptotically stable under the sliding mode controller designed in this paper even though the premise variables are partly unknown.\nExample 2: Consider the nonlinear system as follows: \u23a7 \u23aa\u23aa\u23aa\u23aa\u23a8\n\u23aa\u23aa\u23aa\u23aa\u23a9\nx\u03071 = x1 + x2 + sin(x3) + 0.1x4 + u + w(x1, x2, x3, x4) x\u03072 = x1 \u2212 2x2 x\u03073 = x1 + x2 1x2 \u2212 0.3x3 x\u03074 = sin(x3) \u2212 x4 y = x1 + x3. (75)\nAccording to the T\u2013S fuzzy modeling theory, we can construct the following T\u2013S fuzzy system to represent the nonlinear system (75).\nPlant Rule 11: If f (x) is \u03b811 and g(x) is \u03b821, Then\nx\u0307 = A11x + B(u + w(x)) y = Cx. (76)\nPlant Rule 12: If f (x) is \u03b811 and g(x) is \u03b822, Then\nx\u0307 = A12x + B(u + w(x)) y = Cx. (77)\nPlant Rule 21: If f (x) is \u03b812 and g(x) is \u03b821, Then\nx\u0307 = A21x + B(u + w(x)) y = Cx. (78)\nPlant Rule 22: If f (x) is \u03b812 and g(x) is \u03b822, Then\nx\u0307 = A22x + B(u + w(x)) y = Cx (79)\nwhere\nf (x) = x2 1, g(x) = sin(x3)/x3, \u03b811 = x2 1/a2 1, \u03b812 = 1 \u2212 \u03b811\n\u03b821 = a2(sin(x3) \u2212 x3)/(x3(sin(b) \u2212 b)), \u03b822 = 1 \u2212 \u03b821\nx1 \u2208 [\u2212a1, a1], g(x) \u2208 [gmin, gmax] = [ sin(a2)/a2, 1] x3 \u2208 [\u2212a2, a2], f (x) \u2208 [fmin, fmax] = [0, a2 1]\nA11 =\n\u23a1\n\u23a2 \u23a2 \u23a3 1 1 1 \u22120.1 1 \u22122 0 0 1 a2\n1 \u22120.3 0 0 0 1 \u22121\n\u23a4\n\u23a5 \u23a5 \u23a6, B =\n\u23a1\n\u23a2 \u23a2 \u23a3 1 0 0 0\n\u23a4\n\u23a5 \u23a5 \u23a6",
            "A12 =\n\u23a1\n\u23a2 \u23a2 \u23a3 1 1 sin(a2/a2) \u22120.1 1 \u22122 0 0 1 a2\n1 \u22120.3 0 0 0 sin(a2)/a2 \u22121\n\u23a4\n\u23a5 \u23a5 \u23a6\nA21 =\n\u23a1\n\u23a2 \u23a2 \u23a3 1 1 1 \u22120.1 1 \u22122 0 0 1 0 \u22120.3 0 0 0 1 \u22121\n\u23a4\n\u23a5 \u23a5 \u23a6, CT =\n\u23a1\n\u23a2 \u23a2 \u23a3 1 0 0 0\n\u23a4\n\u23a5 \u23a5 \u23a6\nA22 =\n\u23a1\n\u23a2 \u23a2 \u23a3 1 1 sin(a2/a2) \u22120.1 1 \u22122 0 0 1 0 \u22120.3 0 0 0 sin(a2)/a2 \u22121\n\u23a4\n\u23a5 \u23a5 \u23a6.\nIn this example, the system state x3 is assumed to be unmeasured, that is, the premise variable g(x) is unknown. Thus, the conventional SMC approach cannot be used. By solving (8), we can obtain\nA1 =\n\u23a1\n\u23a2 \u23a2 \u23a3 0.9209 1.0129 0.8325 \u22120.0768 1.0296 \u22122.0010 \u22120.0000 0.0078 0.9255 4.0999 \u22120.3000 \u22120.1117 0.0036 \u22120.0113 0.8325 \u22121.0448\n\u23a4\n\u23a5 \u23a5 \u23a6 (80)\nA2 =\n\u23a1\n\u23a2 \u23a2 \u23a3 0.9547 0.9401 0.8325 \u22120.0538 0.9553 \u22121.9219 \u22120.0000 \u22120.0309 0.9635 \u22120.0749 \u22120.3000 \u22120.0843 0.0825 0.0035 0.8325 \u22120.9709\n\u23a4\n\u23a5 \u23a5 \u23a6.\n(81)\nUsing Corollary 3, we can obtain the control gains K1 = 2.7855 and K2 = 3.3027. Select the parameters of controller as \u03b3 = 0.05, \u03b2\u03041 = 0.15, \u03b2\u03042 = 0.12, \u03b41 = 0.2, and \u03b42 = 0.25. Then, the system state trajectories are plotted in Fig. 5. Figs. 6 and 7 show the SMC signal and sliding mode surface, respectively.\nFrom Figs. 5\u20137, it is obvious that the designed outputfeedback sliding mode controller can stabilize the T\u2013S fuzzy system with case 2. Moreover, different from [5], [9], [12], [16], [19], and [31] and other filter-based results, for the TSFSS with partly immeasurable premise variables, the SMC method proposed in this paper is simple, the controller can be directly obtained, and the observer and filter are unnecessary to be designed.\nFrom Examples 1 and 2, the effectiveness of the proposed SMC design methods in this paper can be adequately verified.\nV. CONCLUSION\nIn this paper, novel sliding mode controller design methods for TSFSS with partly immeasurable premise variables have been derived. Unlike the existing results on observer-based SMC, the SMC design methods in this paper do not require to design an observer when the system states are partly measurable. The TSFSS with partly immeasurable premise variables has been represented by a new TSFSS with totally measurable premise variables. The stability, controllability, observability, regularity, and impulsive behavior of the new TSFSS are same as the original system. Considering different cases of immeasurable premise variables, different sliding mode surfaces have been developed. By analyzing the sliding mode dynamic, the stability conditions have been obtained. Then, the state-feedback and static output-feedback sliding mode controllers have been obtained. Finally, two examples have been given to verify the effectiveness of the proposed methods. In the future, we will further study the application of SMC in TSFSSs.\nREFERENCES\n[1] H. H. Rosenbrock, \u201cStructural properties of linear dynamical systems,\u201d Int. J. Control, vol. 20, no. 2, pp. 191\u2013202, 1974. [2] D. G. Luenberger and A. Arbel, \u201cSingular dynamic Leontief systems,\u201d Econometrica J. Econometric Soc., vol. 45, no. 4, pp. 991\u2013995, 1977. [3] N. Mcclamroch, \u201cSingular systems of differential equations as dynamic models for constrained robot systems,\u201d in Proc. IEEE Int. Conf. Robot. Autom., vol. 3, 1986, pp. 21\u201328. [4] T. Taniguchi, K. Tanaka, K. Yamafuji, and H. O. Wang, \u201cFuzzy descriptor systems: Stability analysis and design via LMIs,\u201d in Proc. IEEE Amer. Control Conf., vol. 3, 1999, pp. 1827\u20131831. [5] C. Wu, J. Liu, Y. Xiong, and L. Wu, \u201cObserver-based adaptive faulttolerant tracking control of nonlinear nonstrict-feedback systems,\u201d IEEE Trans. Neural Netw. Learn. Syst., vol. 29, no. 7, pp. 3022\u20133033, Jul. 2018.",
            "[6] Y. Pan and G.-H. Yang, \u201cEvent-triggered fuzzy control for nonlinear networked control systems,\u201d Fuzzy Sets Syst., vol. 329, pp. 91\u2013107, Dec. 2017. [7] C. Wu, J. Liu, X. Jing, H. Li, and L. Wu, \u201cAdaptive fuzzy control for nonlinear networked control systems,\u201d IEEE Trans. Syst., Man, Cybern., Syst., vol. 47, no. 8, pp. 2420\u20132430, Aug. 2017. [8] Y. Pan and G.-H. Yang, \u201cEvent-triggered fault detection filter design for nonlinear networked systems,\u201d IEEE Trans. Syst., Man, Cybern., Syst., vol. 48, no. 11, pp. 1851\u20131862, Nov. 2018. [9] L. Wang, M. V. Basin, H. Li, and R. Lu, \u201cObserver-based composite adaptive fuzzy control for nonstrict-feedback systems with actuator failures,\u201d IEEE Trans. Fuzzy Syst., vol. 26, no. 4, pp. 2336\u20132347, Aug. 2018. [10] Y. Pan and G.-H. Yang, \u201cNovel event-triggered filter design for nonlinear networked control systems,\u201d J. Frankl. Inst., vol. 355, no. 3, pp. 1259\u20131277, 2018. [11] L. Wang, H. Li, Q. Zhou, and R. Lu, \u201cAdaptive fuzzy control for nonstrict feedback systems with unmodeled dynamics and fuzzy dead zone via output feedback,\u201d IEEE Trans. Cybern., vol. 47, no. 9, pp. 2400\u20132412, Sep. 2017. [12] H. Ma, Q. Zhou, L. Bai, and H. Liang, \u201cObserver-based adaptive fuzzy fault-tolerant control for stochastic nonstrict-feedback nonlinear systems with input quantization,\u201d IEEE Trans. Syst., Man, Cybern., Syst., vol. 49, no. 2, pp. 287\u2013298, Feb. 2019. [13] Y. Pan and G.-H. Yang, \u201cEvent-based output tracking control for fuzzy networked control systems with network-induced delays,\u201d Appl. Math. Comput., vol. 346, pp. 513\u2013530, Apr. 2019. [14] S. Xu, B. Song, J. Lu, and J. Lam, \u201cRobust stability of uncertain discrete-time singular fuzzy systems,\u201d Fuzzy Sets Syst., vol. 158, no. 20, pp. 2306\u20132316, 2007. [15] M. Chadli, H. R. Karimi, and P. Shi, \u201cOn stability and stabilization of singular uncertain Takagi\u2013Sugeno fuzzy systems,\u201d J. Frankl. Inst., vol. 351, no. 3, pp. 1453\u20131463, 2014. [16] Y. Zhang, Z. Jin, and Q. Zhang, \u201cObserver design for a class of T\u2013S fuzzy singular systems,\u201d Adv. Differ. Equ., vol. 2017, no. 1, pp. 1\u201319, Dec. 2017. [17] J. Wang, S. Ma, and C. Zhang, \u201cResilient estimation for T\u2013S fuzzy descriptor systems with semi-Markov jumps and time-varying delay,\u201d Inf. Sci., vols. 430\u2013431, pp. 104\u2013126, Mar. 2018. [18] H. H. Choi, \u201cLMI-based sliding surface design for integral sliding mode control of mismatched uncertain systems,\u201d IEEE Trans. Autom. Control, vol. 52, no. 4, pp. 736\u2013742, Apr. 2007. [19] B. Bandyopadhyay, P. S. Gandhi, and S. Kurode, \u201cSliding mode observer based sliding mode controller for slosh-free motion through PID scheme,\u201d IEEE Trans. Ind. Electron., vol. 56, no. 9, pp. 3432\u20133442, Sep. 2009. [20] H. Wang, Y. Jing, C. Yu, S. Wang, and J. Gao, \u201cGuaranteed cost sliding mode control for discrete-time looper systems in hot strip finishing mills,\u201d J. Control Decis., vol. 3, no. 2, pp. 151\u2013164, 2016. [21] J. Zhao, S. Jiang, F. Xie, X. Wang, and Z. Li, \u201cAdaptive dynamic sliding mode control for space manipulator with external disturbance,\u201d J. Control Decis., p. 16, May 2018. doi: 10.1080/23307706.2018.1487807. [22] C. Wu, Z. Hu, J. Liu, and L. Wu, \u201cSecure estimation for cyber-physical systems via sliding mode,\u201d IEEE Trans. Cybern., vol. 48, no. 12, pp. 3420\u20133431, Dec. 2018. [23] G. Tarcha\u0142a and T. Or\u0142owska-Kowalska, \u201cEquivalent-signal-based sliding mode speed MRAS-type estimator for induction motor drive stable in the regenerating mode,\u201d IEEE Trans. Ind. Electron., vol. 65, no. 9, pp. 6936\u20136947, Sep. 2018. [24] Y. Li, C. Tang, S. Peeta, and Y. Wang, \u201cIntegral-sliding-mode braking control for a connected vehicle platoon: Theory and application,\u201d IEEE Trans. Ind. Electron., vol. 66, no. 6, pp. 4618\u20134628, Jun. 2019. doi: 10.1109/TIE.2018.2864708. [25] L. Wu and W. X. Zheng, \u201cPassivity-based sliding mode control of uncertain singular time-delay systems,\u201d Automatica, vol. 45, no. 9, pp. 2120\u20132127, 2009. [26] L. Wu and D. W. C. Ho, \u201cSliding mode control of singular stochastic hybrid systems,\u201d Automatica, vol. 46, no. 4, pp. 779\u2013783, 2010. [27] X. Sun and Q. Zhang, \u201cAdmissibility analysis for interval type-2 fuzzy descriptor systems based on sliding mode control,\u201d IEEE Trans. Cybern., to be published. doi: 10.1109/TCYB.2018.2837890. [28] J. Li and Q. Zhang, \u201cA linear switching function approach to sliding mode control and observation of descriptor systems,\u201d Automatica, vol. 95, pp. 112\u2013121, Sep. 2018. [29] X. Sun and Q. Zhang, \u201cObserver-based adaptive sliding mode control for T\u2013S fuzzy singular systems,\u201d IEEE Trans. Syst., Man, Cybern., Syst., to be published. doi: 10.1109/TSMC.2018.2852957. [30] J. Li, Q. Zhang, X.-G. Yan, and S. K. Spurgeon, \u201cRobust stabilization of T\u2013S fuzzy stochastic descriptor systems via integral sliding modes,\u201d IEEE Trans. Cybern., vol. 48, no. 9, pp. 2736\u20132749, Sep. 2018. [31] J. Dong and G.-H. Yang, \u201cObserver-based output feedback control for discrete-time T\u2013S fuzzy systems with partly immeasurable premise variables,\u201d IEEE Trans. Syst., Man, Cybern., Syst., vol. 47, no. 1, pp. 98\u2013110, Jan. 2017. [32] J. Dong and Y. Fu, \u201cA design method for T\u2013S fuzzy systems with partly immeasurable premise variables subject to actuator saturation,\u201d Neurocomputing, vol. 225, pp. 164\u2013173, Feb. 2017. [33] J. Dong and G.-H. Yang, \u201cH\u221e filtering for continuous-time T\u2013S fuzzy systems with partly immeasurable premise variables,\u201d IEEE Trans. Syst., Man, Cybern., Syst., vol. 47, no. 8, pp. 1931\u20131940, Aug. 2017. [34] J. Dong and S. Wang, \u201cRobust H\u221e tracking control design for T\u2013S fuzzy systems with partly immeasurable premise variables,\u201d J. Frankl. Inst., vol. 354, no. 10, pp. 3919\u20133944, 2017. [35] Y. Wu and J. Dong, \u201cFault detection for T\u2013S fuzzy systems with partly unmeasurable premise variables,\u201d Fuzzy Sets Syst., vol. 338, pp. 136\u2013156, May 2018. [36] D. W. C. Ho and Y. Niu, \u201cRobust fuzzy design for nonlinear uncertain stochastic systems via sliding-mode control,\u201d IEEE Trans. Fuzzy Syst., vol. 15, no. 3, pp. 350\u2013358, Jun. 2007. [37] Z. Feng and P. Shi, \u201cTwo equivalent sets: Application to singular systems,\u201d Automatica, vol. 77, pp. 198\u2013205, Mar. 2017. [38] B. Jiang, H. R. Karimi, Y. Kao, and C. Gao, \u201cTakagi\u2013Sugeno modelbased sliding mode observer design for finite-time synthesis of semiMarkovian jump systems,\u201d IEEE Trans. Syst., Man, Cybern., Syst., to be published. doi: 10.1109/TSMC.2018.2846656. [39] M. A. Khanesar, O. Kaynak, S. Yin, and H. Gao, \u201cAdaptive indirect fuzzy sliding mode controller for networked control systems subject to time-varying network-induced time delay,\u201d IEEE Trans. Fuzzy Syst., vol. 23, no. 1, pp. 205\u2013214, Feb. 2015. Xingjian Sun received the B.S. degree in mathematics and the M.S. degree in applied mathematics from Bohai University, Jinzhou, China, in 2012 and 2015, respectively. He is currently pursuing the Ph.D. degree in control theory and control engineering with Northeastern University, Shenyang, China. His current research interests include Takagi\u2013 Sugeno fuzzy systems, descriptor systems, sliding mode control, and interval type-2 fuzzy systems. Qingling Zhang received the B.S. and M.S. degrees in mathematics and the Ph.D. degree in automatic control from Northeastern University, Shenyang, China, in 1982, 1986, and 1995, respectively. He was a Post-Doctoral Research Fellow with the Automatic Control Department, Northwestern Polytechnical University, Xi\u2019an, China, from 1997 to 1999. He was a Professor and the Dean of the College of Science, Northeastern University from 1997 to 2006. He visited Hong Kong University, Hong Kong; Seoul University, Seoul, South Korea; Alberta University, Edmonton, AB, Canada; Lakehead University, Thunder Bay, ON, Canada; Sydney University, Sydney, NSW, Australia; Western Australia University, Perth, WA, Australia; Windsor University, Windsor, ON, Australia; Hong Kong Polytechnic University, Hong Kong; and Kent University, Canterbury, U.K., as a Research Associate, a Research Fellow, a Senior Research Fellow, and a Visiting Professor. He has published 16 books and over 600 papers about control theory and applications. Prof. Zhang was a recipient of 14 prizes from central and local governments for his research, and the Golden Scholarship from Australia in 2000. He is currently the Vice Chairman of the Chinese Biomathematics Association, a member of the Technical Committee on Control Theory of the Chinese Association of Automation and the Chinese Association of Mathematics, and the Chairman of the Mathematics Association of Liaoning Province. He was a member of the University Teaching Advisory Committee of the National Ministry of Education."
        ]
    },
    {
        "image_filename": "designv11_13_0002146_j.mechmachtheory.2014.04.012-Figure19-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002146_j.mechmachtheory.2014.04.012-Figure19-1.png",
        "caption": "Fig. 19. Simulated tooth contact ellipses and contact points with ideal gear assembly.",
        "texts": [],
        "surrounding_texts": [
            "In this example, we investigate and compare the contact ellipses and transmission errors of work gear pairs shaved by the proposed double-crowning method with an auxiliary crowning mechanism. The work gear pairs again comprise a double-crowned pinion 2 generated with shaving cutter assembly errors (i.e., cutter center distance error \u0394Es, vertical axial error \u0394vs, and horizontal axial error \u0394hs) and a standard helical gear 3. We also investigate the effects of work gear pair assembly errors (i.e., gear center distance error\u0394Eg, vertical misaligned angle\u0394vg, and horizontalmisaligned angle\u0394hg) on the tooth contact ellipses, contact points on gear 3, and transmission errors in themeshing of work gear pairs. The basic parameters for the pinion 2, gear 3, and shaving cutter are the same as those listed in Table 1. To appraise and compare the contact point dislocation in the meshing of work gear pairs, we propose two indexes: one for appraising center contact point dislocation (designated Lc) and the other for assessing contact point distribution in the longitudinal direction (designated Ld): Case 1: The work gear pairs comprise gear 3 and pinion 2 generated under shaving cutter center distance error \u0394Es = 0.1 mm. Besides, thework gear pairs have assembly errors of gear center distance error\u0394Eg = 1 mm, verticalmisaligned angle\u0394vg = 0.1\u00b0, and horizontal misaligned angle \u0394hg = 0.1\u00b0. The tooth contact ellipses, contact points, and transmission errors of the gear pairs are simulated and shown in Figs. 11\u201318, respectively, which clearly show thatwork gear pair assembly errors have a notable effect on center contact point dislocation. The horizontal misaligned angle, particularly, increases dislocation from 0.78 mm (Fig. 11) to 4.72 mm (Fig. 14). Such errors, however, have little effect on contact point distributions in the longitudinal direction, changing their values onlywithin a small range (Ld = 6.28 mm,\u2026, 6.42 mm), and themaximumcontact point distribution in this direction Ld = 6.42 mm is also much smaller than the face width of gear 3 (Fw3 = 25.4 mm). Hence, as Figs. 11\u201314 show, the proposed double-crowned work gear eliminates the inducement of edge contact on gear 3. In this case, the transmission errors of the proposed gear set are in fact a parabolic function, with a negligible maximum magnitude \u0394\u03d53g = 19.0 arcsec (Figs. 15\u201318). Case 2: In this case, the work gear pairs comprise gear 3 and pinion 2 generated under vertical axial error \u0394vs = 0.05\u00b0. The work gear pairs also have the same assembly errors as those of Case 1. The tooth contact ellipses, contact points, and transmission errors of the gear pairs are simulated and shown in Figs. 19\u201326, respectively. Here, the contact point distributions in the longitudinal direction also change little, but the center contact point dislocation under all gear assembly conditions increases drastically, with the horizontal misaligned angle having the most effect and inducing the risk of edge contact with gear 3 (Figs. 19\u201322). In this case also, the work gear transmission errors for this gear pair are actually a parabolic function, with a negligible maximum magnitude of 21.0 arcsec (Figs. 23\u201326). Case 3: Thework gear pair in this case ismade up of gear 3 and pinion 2 generated under horizontal axial error\u0394hs = 0.05\u00b0with the same gear pair assembly errors as those of Cases 1 and 2. The simulated tooth contact ellipses, contact points, and transmission errors are shown in Figs. 27\u201334, respectively. Here, the center contact point dislocation increases greatly under all gear assembly conditions but is particularly affected by the horizontalmisaligned angle. Themaximumcontact point distribution (Ld = 6.40 mm), however, is much smaller than the face width of gear 3 (Fw3 = 25.4 mm), meaning that, as Figs. 27\u201330 indicate, when the double-crowned pinion 2 is generated under a horizontal axial error of\u0394hs = 0.05\u00b0, edge contact on gear 3 can be avoided. Once again, thework gear transmission errors for this gear pair are in parabolic functions, having a small maximum magnitude of less than 22.0 arcsec (Figs. 31\u201334)."
        ]
    },
    {
        "image_filename": "designv11_13_0000801_tmag.2015.2439043-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000801_tmag.2015.2439043-Figure8-1.png",
        "caption": "Fig. 8. Structure of the novel generator. (a) Diagrammatic sketch. (b) Photo.",
        "texts": [
            " Thus, we obtain 2 2 I 2 0 I II2 2 II 2 0 II III2 2 III 0 17 2 18 17 9 99 18 27 19 1919 2 54 3 l l l l l N NL R R NL L RN NL R R NL L RN NL R R \u23a7 = +\u23aa \u23a7\u23aa = +\u23aa\u23aa\u23aa \u23aa= + \u21d2\u23a8 \u23a8 \u23aa \u23aa = +\u23aa \u23aa \u23a9\u23aa = + \u23aa\u23a9 (5) where LI, LII, and LIII are the inductances of type I, type II, and type III respectively. Equation (5) indicates that concentrating several dispersed coils into one coil is conducive to increasing per phase self-inductance value, and this is the method we use in this paper for increasing the inductances of winding 1. As shown in Fig. 8 (a), the novel generator is a PM synchronous generator with two sets of three-phase winding: winding 1 (phases A, B, and C) and winding 2 (phases U, V, and W). The number of turns of winding 1 is greater than that of winding 2, i.e. N1>N2. In the novel generator, winding 1 employs the concentrated-coil-arrangement, which thereby leads to a high value of per phase self-inductance of winding 1. Per phase self-inductance value of winding 2 should be small because winding 2 must output large amounts of power in order for the total input power curve to approach the MPP curve (see Fig. 5). Therefore, in the novel generator, winding 2 employs the dispersed-coil-arrangement. Table I shows the values of per phase inductances of winding 1 and winding 2, i.e. L1 and L2. As shown in Fig. 8 (b), a prototype was manufactured to validate this generator system concept. Its rated power and rotational speed are 1000W and 400r/min, respectively. Fig. 9 shows the measured characteristics of the prototype system. It can be observed that the input power curve approaches the MPP curve. Fig. 10 (a) shows the experimental setup. A PM synchronous motor system, comprising an inverter, a singlechip microcomputer, and a PM synchronous motor, acted as a wind turbine emulator. The microcomputer single-chip calculated a torque command according a measured rotor speed and the wind turbine characteristics, and then sent the torque command to the inverter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002720_978-3-319-24055-8-Figure3.9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002720_978-3-319-24055-8-Figure3.9-1.png",
        "caption": "Fig. 3.9 Screen shot of the animation of the operating deflected shape (ODS) for the key vibration peak, Mode 0 Force Shape, 36 cycles per revolution",
        "texts": [
            " Figure 3.8 shows the response for the different excitation sources, each of which have different frequencies. The frequency excitation at maximum motor speed varies from 400 Hz for imbalance to 19.2 kHz for Mode 6 Force Shape, 48 cycles. Feedback was given to the design team to change the housing design so as to minimise these modes of vibration. This feedback is commonly given in the form of animated Operating Deflected Shapes (ODS) of the vibration. A screen shot of one such animation can be seen in Fig. 3.9. The complex vibratory motion of the end face of the motor housing, at the far right end of the model, was predicted by the simulation of the compete powertrain structure and it is known that this matches the experience of engineers who have worked on previous motor design projects. As well as providing feedback to the design process for the application of ribs in this area, this indicated that the system simulation was matching the test data from previous projects. The housing design was developed for the selected concept, with the structure modified and ribs applied to reflect the feedback from the intermediate simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure6-1.png",
        "caption": "Fig. 6. Tooth clearances due to the shaft deformation.",
        "texts": [
            " On the other hand, any other planet shaft i will be also deformed in the tangential direction due to the twist of the carrier. In this case (type 2) the deflection \u03b4t i - u ( z k ) of the unit u on the position z k of another planet i is also as a function of F eqv by combination of Eqs. (23) , (24) , (27) and (28) with the calculated twist \u03b4B , \u03b4A . The effective displacement of the contact tooth pair due to the loaded deformation can be regarded as the projected distance on the line of action, as shown in Fig. 6 . The influence coefficients f F i-j for both the types are derived based on the relation of the acting load and the deflection. For example, the equation is valid for type 1, f FC1 = \u03b4t j,u ( z k ) cos \u03b1w F j \u2212l \u2212v , (31) or f F i \u2212 j = \u03b4t j,u ( z k ) F eqv . \u00b7 cos 2 \u03b1w , (32) where the relation \u03b4t j,u ( z k ) / F eqv is a constant and can be obtained from the equations mentioned above. Similar relation is also valid for the influence coefficient of type 2. Besides the influence coefficients, the separation distances between two engaged flanks in the proposed LTCA model must be also determined based on the mesh relation of the gear set [29] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001038_asjc.1502-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001038_asjc.1502-Figure2-1.png",
        "caption": "Fig. 2. TRMS working model [1]. [Color figure can be viewed at wileyonlinelibrary.com]",
        "texts": [
            ") below: \u03c41 \u00bc k1 T 11s\u00fe T10 :u1 (6) Moments in the horizontal plane: I2:\u20ac\u03d5 \u00bc M2 MB\u03d5 MR (7) where \u03d5 is azimuth angle, I2 is moment of inertia in the vertical plane, M2 is nonlinear static momentum in the vertical plane, MB\u03d5 is friction forces momentum in the vertical plane, MR is cross reaction momentum in the vertical plane and M2 \u00bc a2:\u03c422 \u00fe b2:\u03c42 (8) MB\u03d5 \u00bc B1\u03d5: _\u03c8 \u00fe B2\u03d5:sign _\u03d5 (9) MR \u00bc kc T0s\u00fe 1\u00f0 \u00de Tps\u00fe 1 :\u03c41 (10) Tail motor input voltage (control) to output torque relation is given by the T.F., \u03c42 \u00bc k2 T 21s\u00fe T 20 :u2 (11) where u1 and u2 are the control inputs to the main motor and tail motor, respectively. Different vectors used in equations 1\u201311 are illustrated in Fig. 2 [1]. The cross coupling effect of the rotors is illustrated in the block diagram of Fig. 3. This diagram may be obtained from the equations 1\u201311. GMR is the main rotor T.F. from control input (\u00b118volts) to output as motor torque (maximum continuous torque, 14.8 mN-m). At this point, it should be made clear that the motor input (\u00b118volts) is different from the saturation input (\u00b12.5volts) to be given to the SCSI adapter. Later versions of TRMS are available with built-in saturation blocks of \u00b12.5volts in MATLAB"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001893_j.triboint.2015.08.007-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001893_j.triboint.2015.08.007-Figure1-1.png",
        "caption": "Fig. 1. (a) The two PU samples used in the experiments. The mass of both samples was 425 g. (b) Sliding tests and contact area measurements at an angle of inclination of \u03b8. (c) Measurement method of sliding distance. (d) Measurement method of contact area.",
        "texts": [
            " The oil layer was applied to avoid an increase in the contact area during the sliding process due to wear; an increase in the contact area complicates the sliding acceleration and deceleration in a single sliding process [25]. To simplify the sliding behavior, the sliding velocity and contact area of the PU samples were measured on an oiled PMMA surface and characterized as a function of the angle of inclination. The relationships between these factors were identified to investigate the effect of contact area on sliding velocity. In addition, the present study proposed an analytical model to describe these observations, which is based on Couette flow [26] with no pressure gradient. Fig. 1(a) shows the two test samples used in the experiments. Three PU rubber specimens with a smooth convex surface (Iteck Co., Japan) were attached tightly to a steel disk (with an outer diameter of 90 mm, an inner diameter of 20 mm and a thickness of 9 mm). Two PU specimens with different convex surfaces were used to vary the contact area: Sample 1 was 8 mm in diameter and 2.5 mm thick at the center, and Sample 2 was 9 mm in diameter and 3.5 mm thick at the center. The mass of both samples was 425 g. The contact surfaces were wiped with cotton swabs to remove dust prior to each experiment. Contents lists available at ScienceDirect journal homepage: www.elsevier.com/locate/triboint Tribology International http://dx.doi.org/10.1016/j.triboint.2015.08.007 0301-679X/& 2015 Elsevier Ltd. All rights reserved. n Tel.: \u00fe81 92 583 7761; fax: \u00fe81 92 593 3947. E-mail address: k.arakaw@riam.kyushu-u.ac.jp Fig. 1(b) shows the sliding test and the contact area measurement at an angle of inclination of \u03b8. A smooth transparent PMMA plate that was 5 mm thick and 0.5 1 m2 was mounted on a rigid wooden frame to avoid bending and torsion of the plate. To determine the sliding distance, scale marks were glued to the reverse side of the plate; the center of the sample was marked as shown in Fig. 1(c). The PMMA surface was sprayed with silicone oil, and wiped with tissue paper to leave a thin layer of oil. A nonsolvent-type silicone oil spray (Prostaff Co., Japan) was used, to avoid chemical damage to the PMMA and PU rubber samples. The sliding behavior was recorded using a video camera (Handycam HDR-PJ 760V, Sony Corp.), and the sliding distances were determined from still images using an image converter (PlayMemories, Sony Corp.). The PMMA plate was inclined at angles in the range 20o\u03b8o451. The experiment was conducted at room temperature (i.e., 22 1C) and the relative humidity was about 40%. The sliding tests were performed at least three times for each angle, and the result with the median velocity of the three was used as representative. The sliding direction of the samples is indicated in Fig. 1 (a). To minimize the influence of the surface change on velocity, the trajectories of the three rubbers did not cross each other in a single experiment. The contact area was measured under static conditions, as shown in Fig. 1(b). The sample was mounted on an inclined PMMA plate with a point force. The contact surface was illuminated using diffused light from the reverse side of the PMMA plate by inserting a small 1-mm-thick glass plate that was 12 25 mm2. We recorded a dark spot due to diffuse reflection at a magnification Nomenclature A the contact area ac the diameter of the contact area c c\u00bc\u03b3 A2 F the frictional force Fd the dynamic frictional force g the acceleration due to gravity h the thickness of oil layer k k\u00bcmg(sin\u03b8 sin\u03b8c) L the sliding length m the mass of sliding sample N the contact force P the point force t time v the sliding velocity vn the average sliding velocity Greek symbols \u03b3 \u03b3\u00bc\u03b7\u29f8hA \u03b7 the viscosity of oil layer \u03b8 the angle of inclination \u03b8c the critical angle between dynamic and static friction \u03bc the coefficient of sliding friction \u03c4 the shear stress of 50 , using a digital microscope (MJ-302, Sato Shouji Inc., Japan), as shown in Fig. 1(d). The contact area was then characterized based on the diameter of the dark spot. Fig. 2(a) shows the relationship between the sliding length L and time t at \u03b8\u00bc301 (also shown are curves calculated from the analytical model, which will be discussed later). With two of the samples, L increased almost linearly with t, except during the early stages of sliding, due to acceleration. The gradient of L as a function of t was smaller for Sample 2 than for Sample 1 at a given t. The average velocities determined using linear regression were 0",
            " 2(b) shows L as a function of t for two samples at \u03b8\u00bc401. The gradients of these curves for both samples increased as \u03b8 increased from 301, and the gradient for Sample 2 was smaller than that for Sample 1 for a given t. The average velocities determined using linear regression were 0.46 m/s for Sample 1, and 0.17 m/s for Sample 2. The linear relationship between L and t differed significantly from the expected result calculated using F\u00bc\u03bcN, where F is the frictional force, \u03bc is the coefficient of sliding friction, and N is the contact force (see Fig. 1(b)). Here, L is proportional to the square of time (i.e., L\u00bc1/2 (g sin\u03b8 \u03bc N/m) t2, wherem is the mass of the sample, and g is the acceleration due to gravity). These results suggest that the dynamic frictional force is not constant during sliding, and is related to the contact area and the sliding velocity. Fig. 3(a) shows the average velocity vn as a function of \u03b8, where vn was determined using linear regression. For both samples, vn increased as a function of \u03b8, and was smaller for Sample 2 than Sample 1 for a given \u03b8",
            " For both samples, ac decreased gradually as a function of \u03b8 due to the decrease in the contact force, and was larger for Sample 2 than for Sample 1 for a given \u03b8. The mean of the ratio was ac2/ac1 1.25 for 20o\u03b8o451. This suggests that sliding velocity is related to the square of the contact area, since similar values were determined for the two ratios, i.e., v1* /v2* 2.94 and (A2/A1)2\u00bc(ac2/ac1)4 2.44. This relationship was also suggested for non-lubricated sliding friction of PTFE on glass [25]. Fig. 1(b) also illustrates an analytical model to describe the dynamic sliding behavior, which was based on the following assumptions. First, the dynamic frictional force is given by Fd\u00bc\u03c4A, where \u03c4 is the shear stress acting on the interface between the PU rubber and silicone oil layer. Second, the Couette flow [26] with no pressure gradient is used to describe the shear stress, i.e., \u03c4\u00bc\u03b7v/h, where \u03b7 is the viscosity and h is the thickness of the oil layer. The dynamic frictional force Fd is therefore as follows: Fd \u00bc cv\u00bc \u03b3A2v; \u00f01\u00de where c\u00bc\u03b3A2 and \u03b3\u00bc\u03b7\u29f8(hA), which is a parameter related to the ratio between the viscosity and the volume of the oil layer, and is important in understanding the dynamics of sliding friction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002968_s00170-017-0058-y-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002968_s00170-017-0058-y-Figure4-1.png",
        "caption": "Fig. 4 The geometric model without BD",
        "texts": [
            " Thus, the appropriate welding current for thick Invar alloy ranges from 200 to 250 A. Correspondingly, the height and width of weld bead vary from 3.3 to 4.2 mm and 4.6 to 6.0 mm respectively, as shown in Table 2. Accordingly, the special procedure for multi-layer andmultipass MIG welding of Invar alloy plate with a thickness of 19.05 mm was designed as four layers and ten passes where V groove of 60\u00b0, root face of 1 mm thick, space between two half plates of 1.5 mm was selected. The result obtained from geometry modeling with a scale of 1:1 is shown in Fig. 4 where the size of the model is 100 mm \u00d7 100 mm \u00d7 19.05 mm. In this study, the sequence of four-layer and ten-pass welding of Invar alloy is shown in Fig. 4a. During the welding process, the temperature distribution in welding beam and heat-affected zone is highly non-linear due to the effect of heat resource while the temperature away from welding beam transits smoothly. However, the accuracy and the efficiency of simulation depend upon the intensity of the mesh in FE analysis, which means that the sparse grids lead to low simulation accuracy with high efficiency while the fine mesh gives rise to the contrary results. To strike a balance between efficiency and accuracy, the mesh in weld and heataffected zone must be intensive to ensure simulation accuracy, but the mesh away from the welding beam can be sparse properly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001494_01691864.2018.1493397-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001494_01691864.2018.1493397-Figure3-1.png",
        "caption": "Figure 3. Symbols of three-DOF planar PWDR using two wires and an active balancer.",
        "texts": [
            " These two tensions, \u03b1ctr and \u03b1, are different because the resultant \u03b1 is affected by the actuators\u2019 dynamics in Equation (5). For vectors \u03b1ctr and \u03b1, the direction pulled by the wire is defined as positive. Therefore, the positive direction of the wire length qi is the opposite of that of wire tension \u03b1i because the positive tensile force reduces the wire\u2019s length. In this paper, the end-point of each wire on the actuators are considered constant, and the forms of wires are represented as straight lines. As shown in Figure 3, the wire lengths qi (i = 1, 2) are provided by qi = ||XPi(x, y, \u03b8) \u2212 XAi||, (6) where \u2016 \u00b7 \u2016 denotes the Euclid norm; XPi(x, y, \u03b8) is the position vector of the fixed point of eachwire on the plate; XAi is the position vector of the exit point of the wire on the actuator, which is approximated as a constant. As shown in Figure 3, the force and moment vector f w = (fx, fy, \u03c4)T of the plate, which are generated by the resultant wire tension \u03b1, is f w = W\u03b1. (7) Matrix W \u2208 3\u00d72 has the column vectors wi \u2208 3\u00d71 as W = (w1,w2). Vector wi is written as wi = [ pi rwi \u00d7 pi ] , (8) pi \u2208 2\u00d71 is the unit vector (\u2016pi\u2016 = 1), which indicates the direction of the wire as shown in Figure 3; rwi is the vector from the control point x to the wire-connected point. The symbol \u2018\u00d7\u2019 represents the cross product. Substituting Equation (5) into Equation (7), the following equation is obtainable: f w = W ( \u03b1ctr + Aq\u0308 + Bq\u0307 ) . (9) By differentiating the inverse-kinematics in Equation (6), the following relation is derived [1]: q\u0307 = \u2212WTx\u0307, (10) where x = (x, y, \u03b8)T. Furthermore, differentiating Equation (10) yields q\u0308 = \u2212W\u0307Tx\u0307 \u2212 WTx\u0308. (11) Substituting Equations (10) and (11) into Equation (9) yields f w = W ( \u03b1ctr \u2212 AWTx\u0308 \u2212 (AW\u0307T + BWT)x\u0307 ) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002865_j.ijheatmasstransfer.2016.10.057-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002865_j.ijheatmasstransfer.2016.10.057-Figure6-1.png",
        "caption": "Fig. 6. Experimental setup with the welding optics, a coaxial integrated thermography sensor camera and the work piece.",
        "texts": [
            "2016 The experimental benefit of the preceding analytical framework is now exemplified for a real laser welding process, for which we check the validity of the analytically predicted scaling behavior of the geometric quantities characterizing the surface temperature with the feed rate and the source power. Bead-on-plate keyhole weldings without shielding gas with 8 mm thick mild steel plates (assumed material properties a \u00bc 6:04 10 6 m2=s; k \u00bc 33:6 W=m K; TM \u00bc 1500 K [26]) were performed with a BEO D70 optics of Trumpf with a magnification factor of 1:4 : 1 resulting in a focal spot of 280 lm. The experimental setup is illustrated in Fig. 6. The process parameters were varied in a range of P \u00bc 2 6 kW and v \u00bc 2 6 m=min. The experimental values imply large Pe-numbers Pe 80 . . .240 and length scales in the range of l0 0:05 . . .0:2 mm. The near-field region is therefore divided in an inner and an outer area (see Section 3.5). Under these conditions the dimensions of the melt pools are much larger than the length scale and much smaller than the radius of the nearfield region (rN 60 . . .140 mm for e \u00bc 1%). Hence, the formulas of the outer near-field regionN O can be used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000704_sas.2012.6166307-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000704_sas.2012.6166307-Figure1-1.png",
        "caption": "Fig. 1. Illustration of a basic reference measurement setup.",
        "texts": [
            " Moreover, we show that the well-known zero-crossing frequency demodulation method can represent a valid solution for retrieving the vibrational signal, and propose a measurement setup using more than one probe to increase the measurement accuracy independent of the regularity of the cogwheel shape profile. Most of the results presented in this paper were validated by means of numerical simulations and experimental measurements made at the GE Oil & Gas Facility in Florence, Italy. We refer to an arbitrarily shaped cogwheel (Fig. 1) positioned in front of a sensing probe. The measurement device in Fig. 1 is devised for measuring the variations of the shaft angular velocity \u03c9(t) with respect to its average value. From an operational point of view, the average angular velocity is calculated referring to a finite time-window of length W , i.e., \u2206\u03c9(t) = \u03c9(t) \u2212 1 W \u222b t t\u2212W \u03c9(\u03c4)d\u03c4. (1) In this paper we manage the angular vibrations as small perturbations of a constant rotating shaft with angular velocity \u03c90, i.e., \u03c9(t) = \u03c90 + \u03c9\u0303(t) = \u03c90 + \u2206\u03c9maxm(t), where the term \u03c9\u0303 takes into account the angular vibrations, \u2206\u03c9max is the maximum level of |\u03c9\u0303| and m(t) is nondimensional, with |m(t)| \u2264 1, t \u2265 0 and zero mean value"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000600_0954406213477777-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000600_0954406213477777-Figure6-1.png",
        "caption": "Figure 6. An accelerometer mounted on the gearbox.",
        "texts": [
            " In the MFS experimental setup, three-phase induction motor is mounted to the rotor that is connected to the gear box through a pulley and belt mechanism. The gear box and its assembly are illustrated in Figure 4. In the study of faults in gears, three different types of faulty pinion gears namely the chipped tooth (CT), missing tooth (MT) and worn tooth (WT) along with normal gear (or no defect i.e. ND) were used (illustrated in Figure 5). The real time data in time domain were measured using a tri-axial accelerometer (sensitivity: x-axis 100.3mV/g, y-axis 100.7mV/g, z-axis 101.4mV/g) mounted on the top of the gearbox (illustrated in Figure 6) and the data acquisition hardware. Measurements were taken for the rotational speed of 10 to 30Hz in intervals of 2.5Hz for each of four conditions. For each measurement set, 300 Table 2. Optimization fitness function and design parameters. Fitness function Design parameters Bounds C-SVC with RBF kernel Maximize f(x)\u00bc (number of correctly predicted data/total number of testing data) 100% X \u00bc C T For : 0\u20131 For C: 0\u20131.5 -SVC with RBF kernel Maximize f(x)\u00bc (number of correctly predicted data/total number of testing data) 100% X \u00bc T For : 0\u20131 For : 0\u20131"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000683_icra.2013.6631258-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000683_icra.2013.6631258-Figure13-1.png",
        "caption": "Fig. 13. Motion trajectory of neck.",
        "texts": [
            " In this simulation, the humanoid walked in accordance with the following path: (I) \u2192 (II) \u2192 (IV) \u2192 (VI) \u2192 (VI\u2032) \u2192 (VII) \u2192 (I) \u2192 \u00b7 \u00b7 \u00b7 in Fig. 10 as \u201cEmerged gait.\u201d This transition was selected among all possible transient in Fig. 10 by the closed loop dynamics. Figure 12 depicts trajectory of neck (origin of link-18) in 3-D space representation, excluding a trajectory in transient state corresponding to 0\u201320 steps, this means that the walking motion converged into limit cycle after 21 walking steps. Then, the trajectories that are separated into two plane (x-y plane and y-z plane) from initial condition to 200 steps are shown in Fig. 13. We can confirm that motion concerning left side and right side against the dotted line is symmetric, which means the neck and shoulder swung along with y-axis given at Fig. 13, representing rolling motion of upper body. The right graph representing neck\u2019s motion in y-z plane, the neck swayed in sagittal plane forward and backward with height varying by walking states including heel-striking state in it. These figures implies that visual feedback has stabilized the walking including gait\u2019s transition and motions as shown in Fig. 10, including toe-off state, heel-striking, slipping and change of state variables. B. Analyses based on Dynamic Reconfiguration Manipulability In this section, three kinds of lifting proportional-gain are set to be Kp = diag[20, 290, 1100] (Large lifting-gain), Kp = diag[20, 290, 950] (Medium lifting-gain) and Kp = diag[20, 290, 900] (Small lifting-gain) to compare walkings by DRM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003270_s12008-017-0437-5-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003270_s12008-017-0437-5-Figure13-1.png",
        "caption": "Fig. 13 Volumetric deformations of three isotropic and homogeneous materials: a rest state; the strain limit value is b 0.1, c 0.3 and d 0.5; the top row illustrates compressive deformations, whereas the bottom row illustrates tensile deformations",
        "texts": [
            " A material modelled with strain limit value of 0.3 is tested. As shown in Fig. 12, the feedback force decreases asymptotically towards a minimum value. This behaviour is similar to the tissue relaxation response observed from living tissues [45]. Soft tissues are highly complex in terms ofmaterial compositions; therefore, the capability to handle different materials is essential for producing realistic soft tissue deformation. The deformations of three isotropic and homogeneous materials under tension and compression are illustrated in Fig. 13. The three materials are modelled by a volumetric cubic model containing 1331 chain elements with strain limit values of 0.1, 0.3 and 0.5. The deformation of an anisotropic material is achieved with the same volumetric cubic model by setting different values of strain limit in different directions, whereas the deformation of a heterogeneous material is achieved by setting different values of strain limit to chain elements at different regions. A comparison of deformations of isotropic and homogeneous, anisotropic and heterogeneous materials is illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001673_s11837-018-3242-0-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001673_s11837-018-3242-0-Figure1-1.png",
        "caption": "Fig. 1. Cross-sectional view of a 20-W BLDC SPM motor.",
        "texts": [
            " In situ magnetic alignment was performed with the particles aligned in the direction of the applied field and heated to a temperature T (220 C, 238 C, or 256 C) for 15 min. These temperatures and fields were chosen based on in situ alignment studies of anisotropic bonded magnets.21 After alignment, the sample was cooled to 27 C, and magnetic hysteresis loops were measured. This was done sequentially, varying the aligning magnetic field from 0 T to 3 T in steps of 0.25 T. An existing 20-W BLDC surface-mounted permanent magnet (SPM) motor design1 was adapted for bonded magnets in the isotropic condition (unaligned) as shown in Table I. A SPM topology (Fig. 1) was chosen to concentrate the available magnetic flux, which is greatly reduced relative to sintered Nd-Fe-B, but comparable to a ferrite magnet. The performance of the machine was investigated under steady-state condition at the rated speed. The properties of the permanent magnets were varied to those of the anisotropic bonded magnets aligned Khazdozian, Li, Paranthaman, McCall, Kramer, and Nlebedim at 238 C and fields of 0.25 T, 0.5 T, 0.75 T, and 1 T. Only magnets aligned at 238 C were considered, as this was found to be the optimal alignment temperature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003347_1.5021474-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003347_1.5021474-Figure1-1.png",
        "caption": "FIG. 1. Schematic representation of the device filled with CIS/ZnS QDsFLC mixtures.",
        "texts": [
            " The sample cells were fabricated using indium tin oxide (ITO) coated glass plates. The square shape (25 mm2) electrode patterns were achieved on the ITO coated glass plates using photolithographic technique. The planar alignment was obtained by treating the conducting layer with an adhesion promoter and polymer nylon (6/6). After drying the polymer layer, ITO coated glass plates were rubbed unidirectionally. Two ITO coated glass plates were then placed one over another to form a capacitor type cell (Fig. 1). The cell thickness was fixed to 5 lm by placing a Mylar spacer in-between and then sealed with an UV sealant. The empty cells were calibrated using analytical reagent (AR) grade benzene (C6H6) and CCl4 as standard reference for the dielectric study. We used CIS/ZnS core/shell QDs to disperse into the FLC material into different wt. %. The mixtures were prepared by the dispersion of 0.25, 0.5, and 1 wt. % of CIS/ZnS QDs into the pure FLC material. All mixtures were homogenised in an ultrasonic mixer for 1 h at isotropic temperature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001782_012010-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001782_012010-Figure1-1.png",
        "caption": "Fig. 1. Layout of installation for continuous production of powdered cellulose",
        "texts": [
            " Thus, the task of mechanization of the technology of continuous production of powdered cellulose is to combine the method of steam explosion, extraction of lignin and acid hydrolysis of the formed cellulose into a single production cycle. The problem is solved by a series connection of a steam explosion reactor, reactors for intermediate operations and an acid hydrolysis reactor. To move the processed material from one stage of the process to the other, a screw conveyor system is chosen. An installation for the continuous production of powdered cellulose (Figure 1) comprises a sequentially connected premixing device, a loading and feeding unit, a reactor for steam prehydrolysis of wood raw materials with a transport system, a steam feed and release system, a steam explosion 3 reactor, a cellulose fiber acid hydrolysis reactor with a system for supplying and removing acid, a unit for unloading powdered cellulose. The premixing device 1 is filled with wood raw material and water or an acid solution is fed through the valve 4. The components are mixed by a stirrer 2 operating from the drive 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003473_j.engfailanal.2018.05.015-Figure19-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003473_j.engfailanal.2018.05.015-Figure19-1.png",
        "caption": "Fig. 19. Segment A - material S355J2+N.",
        "texts": [
            " Welding of the removed welded joint was carried out with mandatory preheating from the inner surface over previously welded plates. The welded joint was tested using penetrants. After testing, the tube surface needs to be smoothed with a grinder. Removal and welding of the upper and lower tube segment was performed using identical procedures. Reparation was completed by positioning segments A and B on the outer surface of the tube, as shown in Fig. 18. Welding of segments was carried out with mandatory preheating from the inner surface. Geometry of segment A, which corresponds to the angle of 90\u00b0 is shown in Fig. 19. Segment B is similar to segment A and covers the angle of 45\u00b0. The performed welded joints are displayed in Fig. 20. In order to improve strength, i.e., to reduce tube stress at the crack location, new material was added across the entire tube circumference at the location of crack occurrence on the outer surface. An added outside plate is 20mm thick and 300mm wide (Fig. 20). Fig. 21 shows part of the considered structure after repair and reparation. BWE with a repaired boom has been in operation successfully for a number of years"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002441_iros.2015.7354054-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002441_iros.2015.7354054-Figure2-1.png",
        "caption": "Fig. 2. Shown is a scenario where a quadrotor helicopter has uncertainty in both its trajectory estimation and the obstacle location. The a priori variance Pc(\u03c4) + Z for a given obstacle normal n is represented as nT (Pc(\u03c4) + Z)n. The collision point pc is defined as the point along the trajectory where this transformed variance is some distance from the obstacle based on the selected confidence bound p\u0304. The halfspace is defined such that nT p\u0302(\u03c4,\u2206u) > c where c = nTpc.",
        "texts": [
            " Equation (15), given this approximate representation of the uncertainty, is now redefined such that the robot is considered to be collision-free for all time t \u2208 [0, \u03c4 ] if \u2200t \u2208 [0, \u03c4 ] :: R(p\u0302\u22c6(t)) \u2229 (O \u2295 {\u03c3\u0302n}) = \u2205, (19) where \u03c3\u0302 is the distance calculated from the standard deviation and selected confidence bound p\u0304 where \u03c3\u0302 = anT (\u221a Pc(\u03c4) + Z ) n, (20) where a is a scaling factor that corresponds to a Chi-Squared distribution for the given confidence bound p\u0304. Equations (16) and (17) can now be approximated by the following simplified expression: pc = p\u0302\u22c6 ( argmin t\u2208[0,\u03c4 ] {R(p\u0302\u22c6(t)) \u2229 (O \u2295 {\u03c3\u0302n}) \u0338= \u2205} ) , (21) where if the system has no uncertainty \u03c3\u0302 = 0, then Eqs. (19) and (21) are equivalent to the deterministic solution in [10]. Next, given Eqs. (16) and (19), a linear constraint is defined on the position p\u0302(\u03c4,\u2206u) of the robot at time \u03c4 (see Fig. 2) nT p\u0302(\u03c4,\u2206u) > nTpc. (22) Substituting Eq. (11) from Assumption 3, the constraint on the robot\u2019s position in Eq. (22) can be transformed into a constraint on its change in input \u2206u nTJ(\u03c4)\u2206u > nT (pc \u2212 p\u0302\u22c6(\u03c4)). (23) Equation (9) is approximated using Eq. (23) as minimize: \u2206uTR\u2206u (24) subject to: nTJ(\u03c4)\u2206u > nT (pc \u2212 p\u0302\u22c6(\u03c4)), where solving this convex optimization, such as is done by the RVO library in [29], provides a collision free change in input \u2206u with the control input given to the robot as u+\u2206u"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003278_0954406217743271-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003278_0954406217743271-Figure6-1.png",
        "caption": "Figure 6. Pressure p distribution on the larval fish surface at time t/T\u00bc 0.5 (a, b) and t/T\u00bc 0.9 (c, d).",
        "texts": [
            " The anterior and posterior parts of the body generally produce the opposite force in the X direction, leading to a smallCX during the entire duration. In theY direction, the anterior and posterior parts also produce opposite forces. Under similar kinematics, Borazjani simulated the adult sunfish and presented the force production during the entire duration.12 The result shows that the forces in the initial oriented direction and escaping direction share a similar force magnitude, which is consistent with the present observation. Figure 6 shows the pressure p \u00bc p= 0:5 U2 distribution on the larval fish body surface at two typical times at the preparatory stage t=T \u00bc 0:5 and the propulsive stage t=T \u00bc 0:9. At both time points, the posterior part dominates the force production. At t/T\u00bc 0.5, the concave side experiences the high pressure while the convex side experiences the low pressure. At t/ T\u00bc 0.9, on the posterior part, the pushing part generates a large high pressure region while the lee side generates a very low pressure region",
            " Intuitively, whatever the fish intends to capture a prey or escape from a dangerous circumstance, quick acceleration and fast movement are greatly favoured to reduce the time for moving the intended distance. However, many fish decelerate during the quick start and turn case.3,25 Why do the fish undertake this process? A simple model was proposed to answer this question. Figure 8(a) shows the situation at t=T \u00bc 0:5, when a considerable deceleration takes place. From the pressure distribution in Figure 6(a) and (b), the large pressure is generated at the tail, and the force directs to opposite to the escaping direction. The torque is calculated based on the mathematical formula T \u00bc R r0 f dS, where r0 is the position vector with respect to the instantaneous centre of mass. Under this definition, d dx \u00f0 r0i mivi\u00de \u00bc r0i Fext i is satisfied, where vi and Fext i are the velocity and hydrodynamic force in laboratory coordinate system, respectively. We found that the resultant torque on tail is in the clockwise direction (Figure 8(b))"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000250_icaci.2013.6748525-FigureI-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000250_icaci.2013.6748525-FigureI-1.png",
        "caption": "Fig. I. Underwater glider coordinate and variables",
        "texts": [],
        "surrounding_texts": [
            "I. INTRODUCTION\nU NDERWATER gliders comprise \ufffd class of auton\ufffd)mo\ufffds underwater vehicles (AUVs) whIch control theIr attI tude by shifting internal ballast and use changes in buoy ancy as the source of propulsion. An underwater gilder is typically equipped with fixed wings and tails. Unlike conventional AUVs that are battery-powered and propeller driven, the buoyancy-driven gliders travel by gliding upwards and downwards and follow sawtooth-like pattern. The con cept of underwater gliders was conceived by Stommel in 1989 [1]. The idea inspired the design and application of several operational gliders, such as including the SLOCUM glider [2], the Spray glider [3], and the Seaglider [4]. One characteristic of gliders is that they are capable of offering high endurance over long ranges. A well designed gilder can have an expected range above 1500 miles, lasting weeks or months per deployment [5]. The mechanism of gliders results in many useful features for oceanic operations, such as low operating cost, low noise and vibration, and low vulnerability of actuator mechanisms to the harsh effects of seawater. Gliders are attractive for several applications, including remote sensing for oceanography, data collection, environmental monitoring, and military actions [6].\nControl analysis of underwater gliders is an area that has attracted much attention. For example, a nonlinear dynamic model was derived and model-based linear control laws were presented in [7]; a control law was developed by choosing appropriate output variables and using input-output feedback linearization in [8]; a dynamic modelling of the complete multi-body control system and numerical implementation of\ny. Shan and 1. Wang are with School of Control Science and En gineering, Dalian University of Technology, Dalian, Liaoning, China; Z. Yan and 1. Wang are with the Department of Mechanical and Automa tion Engineering, The Chinese University of Hong Kong, Shatin, New Territories, Hong Kong (emails: shanyuan2008@mai1.dlut.edu.cn. {zyan, jwang}@mae.cuhk.edu.hkJ.\n. The work described in the paper was supported by by the NatIonal Natural Science Foundation of China under Grant 61273307; and by the Research Grants Council of the Hong Kong Special Administrative Region, China under Grants CUHK417209E.\n978-1-4673-6343-3/13/$3\\.00 \u00a92013 IEEE 328\na motion control system were described in [9]; a glider coordinated control system (GCCS) using a detailed glider model for prediction and a simple particle model for planning was developed for coordination control of underwater gliders in [10]; control of the spiraling motion of underwater gliders was studied in [II]. Due to its underactuation property and persistent disturbances from ocean environment, controlling an underwater glider is a challenging task. There is always an incentive to further improve the control performances and reliability.\nModel predictive control (MPC) is a powerful control tech nique for optimizing system performances [12][13]. MPC performs real-time optimization based on a system model over a finite horizon of future time. By taking the current state as an initial state, an objective function is optimized at each sampling instant, and the computation is repeated with new state information during the next computation time interval. MPC has been widely applied in real-world applications in past two decades [14]. Applying MPC for underwater glider control could be advantageous. Firstly, MPC adapts to multivariable model naturally. Secondly, MPC takes physical constraints into consideration. Thirdly, MPC iteratively predicts and optimizes system performance so that the environmental changes can be considered.\nThe success of an MPC approach depends on the compu tational efficiency of its optimization method. In this paper, a recurrent neural network is applied for real-time optimization in MPC. There are several reasons for choosing neural net work method: Firstly, neural networks can efficiently solve large scale problems due to the parallel information process ing nature; secondly, neural networks can solve problems with time-vary parameters; thirdly, neural networks can be implemented in hardware. In the past two decades, various recurrent neural network models have been proposed for solving various optimization problems, such as the neural network models presented in [15], [16], [17], [18], [19], [20]. Inspired by the recent success of neurodynamic optimization, some studies on recurrent neural networks based MPC were carried out. For example, a three-layer neural network was applied for constrained MPC in [21]. A dual neural network was applied for the multi-stage optimization problems of generalized predictive control [22]. A two-layer recurrent neural network was applied for solving reformulated mini max optimization problems of robust MPC approaches [23], [24]. The simplified dual network was applied for solving real-time quadratic optimizations of various MPC approaches [25], [26], [27]. Recurrent neural networks greatly improved the computational efficiency.",
            "The structure of the remainder of this paper is as follows. Section II outlines an underwater glider model. In Section III, the MPC problem is formulated as a quadratic programming. In Section IV, the simplified neural network is applied for solving the quadratic programming problem. Finally, section V concludes this paper.\nII. AN UNDERWATER GLIDER MODEL\nTo model the buoyancy and motion dynamics of un derwater gliders in the longitudinal plane, the following assumptions are used throughout the paper:\nAssumption 1: The underwater glider is symmetrical through its lateral and longitudinal body coordinate planes.\nAssumption 2: The mass moment of inertia of the vehicle about the y-axis remains constant.\nAssumption 3: For the longitudinal model, the underwater glider only moves in the vertical x-z plane (see Figure 1) with no yaw moment. So hydrodynamic effects such as side force and induced rudder are neglected in this model.\nAssumption 4: The added mass of accelerating fluid dur ing unsteady motion is negligible.\nThe dynamics of the underwater glider in the longitudinal plane can be described as follows [28]:\ne 'tiL\n. Tnsl Tns3 ihg . V q S111 ex cos ex(-- - --) - --cos ex S111 e Tns3 m'sl Tnsl\nm,g sin2 ex cos2 ex +-cosesinex - D( -- + -- ) rns3 rns3 TrLsI 1 1 +Lsinexcosex(-- - --) , rnsl TrLs3 (Tns3 . 2 m'sl 2 ) mg ( sinexsine\nq -- S111 ex + --cos ex + - , rnsl Tns3 V 'rnsl cos ex cos e ) D . (1 1 ) L ( sin 2 ex\n+ + - S111 ex -- - -- - - --Tns3 V Tnsl Tns3 V Tnsl cos2 ex ) + --- , m's3 1 . - (l'vl - Tnmrg(TmrI cos e + Tmr3 S111 e) Is3 -Tnrbg(TrbI cose + Trb3 sine) + (Tns3\nq,\nUm,\nwhere\nD=(KDO + KDex2)V2, L=(KLO + KLex)V2,\nM=(KMO + KMex + Kqq)V2.\nThe definitions of variables in (1) are listed in Table I. Underwater gliders travel by gliding up-and-down and\nfollow a sawtooth-like pattern in the longitudinal plane. The control objective is to design buoyancy pump rate (Um) and mass moving rate (Umrx) such that the glider follows a desired sawtooth pattern (see Fig. 2).\nIII. MPC FORMULAT ION\nRewrite (1) as an equivalent discrete-time model as follow:\nx(k + 1) = f(x(k)) + g(x(k))u(k), y(k) = Cx(k), (2)\nsubject to constraints:\nUmin ::.; u(k) ::.; Umax, \ufffdUmin ::.; \ufffdu(k) ::.; \ufffdUmax'\nXmin ::.; x(k) ::.; Xmax, Ymin ::.; y(k) ::.; Ymax, (3)\nwhere x = [Vex q e 'tiL Tmrx]T E \ufffd6 is the state vector, U = rUm Umrx]T E \ufffd2 is the input vector, Y = e E \ufffdI is the output vector, C E \ufffdI X \ufffd6 is the output matrix, f and g are nonlinear functions derived from (1), Umin ::.; Umax, \ufffdUmin ::.; \ufffdUmax' Xmin ::.; Xmax, Ymin ::.; Ymax are\nTmrx Umrx, (I) respectively the vectors of lower and upper bounds.",
            "MPC is an iterative optimization based control methodolo gy: at each sampling time k, measure or estimate the current state, then obtain the optimal input vector by solving a real time optimization problem. For model (2), the following cost function is commonly used for MPC design:\nN Nu.-1 J = L Ilr(k + j) -y(k + jlk)II\ufffd+ L II\ufffdu(k + jlk)II\ufffd,\nj=l j=O (4) where r( k + j) denotes the reference trajectory of output signal, y(k + jlk) denotes the predicted output, and \ufffdu(k + jlk) denotes the input increment, where \ufffdu(k+jlk) = u(k+ jlk) -u(k -1 + jlk). N and Nu are prediction horizon (1 :::; N) and control horizon (0 < Nu :::; N), respectively. Q and R are appropriate weighting matrices. 11\u00b711 denotes the Eulidean norm of the corresponding vector. The first term in (4) represents the error between the predicted output and the reference output while the second term considers the control energy. It was shown that there always exists a finite horizon length for which the MPC is stabilizing without the use of a terminal cost or constraint in constrained nonlinear systems [29]. Hence with proper selection of N, Nu, Q, and R, the cost function (4) can also guarantee closed-loop stability.\nAccording to model (2), future state x(k + jlk), j = 1,2, . . . ,N at sampling instant k can be predicted by using the optimal input obtained at previous time instant, i.e., u(k + jlk -1) ,j = 1,2, . . . , Nu. Define the following vectors:\nx(k) = [x(k + 11k) ... x(k + Nlk)]T E 1)t6N, ii(k) = [u(klk) ... u(k + Nu -llk)f E 1)t4Nu., y(k) = [y(k + 11k) ... y(k + Nlk)f E 1)tNm, r(k) = [r(k + 11k) ... r(k + Nlk)f E 1)tNm,\n\ufffdii(k) = [\ufffdu(klk) ... \ufffdu(k + Nu -llk)]T E 1)t4N\". (5)\nThe predicted output y(k) is then expressed in the following form:\ny(k) = O(O\ufffdii(k) + 1 + g), (6)\nwhere\nr g(x(klk - 1)) g(x(k + 11k - 1))\nG=\ng(x(k + N:- 11k - 1)) E lR6NX2Nu.,\nf= [ f(x(klk -1)) ] f(x(k + 11k -1)) f(x(k + N : - 11k -1)) E 1)t6N,\n[ g(x(klk -1))u(k -1) ] g(x(k + 11k -1))u(k -1) g= g(x(k + N - 11\ufffd -1))u(k -1) E 1)t6N.\nHence, the original optImIzation problem (6) subject to constraint (5) becomes\nmm I lr(k) -O(O\ufffdii(k) + 1 + g)ll\ufffd + II\ufffdii(k)II\ufffd, s.t. \ufffdiimin :::; \ufffdii(k) :::; \ufffdiimax'\nwhere\niimin :::; ii(k -1) + i\ufffdii(k) :::; iimax, Xmin :::; 1 + 9 + O\ufffdii(k) :::; Xmax, Ymin :::; 0(1 + 9 + O\ufffdii(k)) :::; Ymax, [1 0 ;] I i= E 1)t2N\"x2Nu.. I\n(7)\nProblem (7) can be rewritten as a time-varying quadratic programming (QP) problem:\nmin s.t.\n\ufffd\ufffdiiTW \ufffdii + cT \ufffdii, l :::; E\ufffdii :::; h,\nwhere the coefficients are:\nW = 2(OTOTQOO + R) E 1)t2Nu.X2N\", C = _20TOTQ(t(k) - Og -01) E 1)t2Nu.,\nE= [-i i -0 0 -co 00 If E 1)t(12Nu.+12N+2Nm)X2Nu.,\nb= -iimin + ii(k -1) iimax -ii(k -1) -Xmin + f + 9 Xmax -1-9\n-Ymin + 01 + Og Ymax - 01-Og\nE 1)t8Nu.+12N+2Nm,\nl = [ \ufffdoo ] E 1)t12Nu.+12N+2Nm, \ufffdUmin h = [ _b ] E 1)t12N, +12N+2Nm. \ufffdumax\n(8)\nThe solution to the QP problem (8) gives optimal control increment vector \ufffdii( k) whose first element \ufffdu( k) can be used to calculate the optimal control input.\nIV. NEURAL NET WORK ApPROACH\nIn [16], based on duality theory, a one-layer recurrent network called the simplified dual network was developed for solving quadratic programming problems. It has been shown good performance and low computational complexity. Its dynamic model can be described as follows: \u2022 State equation\nd z\nE dt = -E\ufffdii + g(E\ufffdii - z ) ,"
        ]
    },
    {
        "image_filename": "designv11_13_0002808_s11071-016-3072-y-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002808_s11071-016-3072-y-Figure1-1.png",
        "caption": "Fig. 1 3D model of the manipulator",
        "texts": [
            " The autoloader includes a transfer device for withdrawing ammunitions from the magazine and inserting withdrawn ammunitions into the gun breech, and a lifting device for alternately moving the transfer device between the withdrawing position and the loading position along the guide track incorporated in the frame. The transfer device, which is mounted on the lifting device, is free to rotate in elevation about an elevation axis on the lifting device. The transfer device is actuated by a pivot motor through two reducers, while the lifting device is driven through a chain drive system, which comprises of driving sprocket, driven sprocket, tension sprocket and the chain interlinking. 2.2 Dynamic model As shown in Fig. 1, oscillations of the tank vehicle on various landforms mainly involve three kinds of movement: (1) shake movement (linear movement along the axis yb); (2) pitch movement (rotational movement about the axis xb); and (3) roll movement (rotational movement about the axis zb),where the Obxbyb is the space coordinate attached to the tank vehicle (which is not shown in the figure). For simplicity, this paper focused on pitch movement only as an initial study. Ignoring the dynamics of the chain driven system, the above manipulator is simplified as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003278_0954406217743271-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003278_0954406217743271-Figure2-1.png",
        "caption": "Figure 2. (a) The non-uniform Cartesian mesh around the larva (only 1 out of every 10 points in each direction is shown). (b) The unstructured triangular surface mesh of the larval fish body.",
        "texts": [
            "20 and Tian et al.21 Then force is calculated by F \u00bc R 1 n dS, where n is surface normal and dS is the infinitesimal area on body surface area . The background pressure, which is picked at the corner of the flow domain, is set as zero. Therefore, the integration of stress over the cross-section of the larval fish leads to zero on force. The fluid\u2013body interaction is implemented using the explicit projection method. A fixed, non-uniform, single-block Cartesian mesh is employed to discretise the fluid domain (Figure 2(a)). The size of the rectangular domain is 5L 5L 2L, with a region near the larval fish body having the maximum resolution 1=250L in all three directions (L is the body length). Increasing the domain size by 1.5 times of the current size in each direction with the same resolution around the fish only leads to the maximum force difference of 1.31% in the Y direction, which means the current domain size is sufficient. The total number of Cartesian grid in the baseline simulation is 18 million (410 360 120). A coarser and a finer case with the finest resolution 1=200L and 1=300L were simulated for the mesh convergence study. The extra simulations produce a maximum of 4.6% and 4.0% difference, respectively, from the baseline mesh in the maximum force in the Y direction. An unstructured three-dimensional surface mesh was implemented to present the surface of the larval fish, as shown in Figure 2(b). The total node and element number are 2492 and 4980, respectively. A surface mesh of 50% finer gave the simulated force less than 4.2% difference in the Y direction. The Courant\u2013 Friedrichs\u2013Lewy (CFL) condition number, which is defined as CFL \u00bc u x \u00fe v y \u00fe w z t \u00f03\u00de is less than 0.5 in the entire domain throughout the simulation, where u, v, w are the velocity components in the x, y, z directions, respectively, and x, y, z and t are maximum spacial resolutions in the three orthogonal directions and time-marching resolution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000335_j.euromechsol.2012.10.011-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000335_j.euromechsol.2012.10.011-Figure3-1.png",
        "caption": "Fig. 3. (a) The elastica passes through the two pads of clamp B. The upper (lower) pad exerts normal force RU (RD) and friction TU (TD) on the elastica. The two pads of clamp B are pre-compressed by force F1. (b) The free body diagram of the upper pad. N is the support reaction in the x-direction exerted by the frame to keep the translational springs staying in the vertical direction.",
        "texts": [
            " After obtaining the elastica deformation, we know the shear force QA at end A and rotation angle qB at end B. We can then find the resultant reactions TB and RB exerted by clamp B in the transverse and longitudinal directions, respectively, from the force balance in the y-direction. We next examine whether the friction exerted by clamp B on the elastica is large enough for the stuck-up assumption to be valid. We consider the situationwhen the two pads of clamp B are precompressed by force F1. The upper pad of clamp B exerts normal force RU and friction TU on the elastica, as shown in Fig. 3(a). Similarly, the lower pad exerts forces RD and TD. The forces in both pads should have the same ratio, m \u00bc TU RU \u00bc TD RD (15) If the ratio m is smaller than mB, then we can be sure that the stuck-up assumption is valid. The next challenge is to find m from the elastica deformation we have solved. The resultants exerted by clamp B on the elastica are TB \u00bc TD \u00fe TU (17) Fig. 3(b) shows the free body diagram of the upper pad of clamp B. The pad is pre-compressed by force F1 in the y-direction. N is the support reaction in the x-direction exerted by the frame to keep the translational springs staying in the vertical direction. From the force balance in the y-direction, and by using Eq. (15), we obtain RU \u00bc F1 cos qB \u00fe msin qB (18) Similarly, from the force balance of the lower pad, we obtain RD \u00bc F1 cos qB msin qB (19) By combining Eqs. (15)e(19), it is found that RB \u00bc F1 2msin qB cos2qB m2sin2qB (20) TB \u00bc F1 2mcos qB cos2qB m2sin2qB (21) From the force balance of the whole beam between A and B in the x-direction, one can eliminate RB and TB to obtain FA \u00bc F1 2m cos2qB m2sin2qB (22) In the case when qB \u00bc 0, Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001273_1077546317742506-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001273_1077546317742506-Figure2-1.png",
        "caption": "Figure 2. Five conditions of sun gears: (a) normal gear; (b) gear with one missing tooth; (c) gear with wear; (d) broken gear; (e) gear with a tooth root crack.",
        "texts": [
            " The signal acquisition system is controlled by epistasis software installed on the computer, for the acquisition of the vibration signals. The detailed specification of the planetary gearbox of the mechanical fault comprehensive simulation bench is shown in Table 2. Because broken teeth, cracks, and wear are common faults in planetary gears, five conditions of sun gears are simulated in this paper. Consistent operating conditions of the simulation bench were ensured during each signal acquisition. Different sun gears are shown in Figure 2: a normal gear, a gear with one missing tooth, a gear with wear, a broken gear, and a gear with a tooth root crack. The motor speed controlled by the software is set as 2400 rev/min. The vibration signals for the five gear conditions are shown in Figure 3. As shown in the five types of vibration signal, the vibration signal of a normal gear has a certain impulse performance, and the impulse component is complex, due to its manufacture and fabrication. When a gear loses a tooth, the impulse component introduced in the meshing process is different from a normal gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000757_j.cirpj.2015.08.005-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000757_j.cirpj.2015.08.005-Figure2-1.png",
        "caption": "Fig. 2. Three magnets group (each group comprised of three individual magnets) located in the master array and slave array.",
        "texts": [
            " Additionally, a formulation is proposed, and experimentally validated, to explain the lateral magnetic force using Bessel function of the first kind. The mobile magnetic clamp has three gripping sets forming, with the contact mechanisms, a kinematic contact system to avoid redundancy and contact ambiguity in planar gripping. Three permanent cylindrical magnets are used at each group to support the axial force generated by the milling operation. Each group has three magnets at the vertices of an equilateral triangle. These three groups of permanent magnets are located on a circumference 120 degrees apart completing the array (Fig. 2). Similar arrangements are adopted both in the master and slave array. A magnet in the master module array has an opposite magnet in the slave module array, forming a pair. The magnets of each pair are nominally coaxial. The clamping contact point is at the centroid of each group where a ball transfer unit is attached (Fig. 3). A magnet attracts the opposite pole and repels the similar pole. Since these permanent magnets are axially magnetized the N-S-N-S orientation is used to get the highest possible attraction force as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000950_6.2017-0011-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000950_6.2017-0011-Figure3-1.png",
        "caption": "Figure 3. Mechanism in straight (top) and swept (bottom) configurations.",
        "texts": [
            " Variable-Sweep Wing Design The starting point for our wing-morphing aircraft was the low-cost flat-foam hobby aircraft shown in Figure 2, which has a wingspan of 42 centimeters and a mass of 37 grams. Several objectives guided the design of the wing sweeping mechanism. First, we sought to make minimal modifications to the original airplane. Second, the mechanism needed to be extremely lightweight to minimize the impact on flight performance. Lastly, we wanted to drive the mechanism with the same small linear servos used to actuate the control surfaces on the airplane. Figure 3 depicts our single-degree-of-freedom modified sliding-crank mechanism design. The red links are attached to the wings, while the yellow link is fixed to the fuselage. The pivot point locations were chosen to provide a 25\u00b0 sweep angle and to allow the ailerons to clear the fuselage. Several prototypes have been constructed from lasercut fiberglass and carbon fiber to provide high stiffness and low weight. A fully assembled airplane with the sweeping mechanism installed is shown in Figure 2. In the remainder of the paper, we analyze the aerodynamic properties of the airplane at the extreme configurations of 0\u00b0 and 25\u00b0 and leave flight experiments with the variable-sweep aircraft to future work"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000021_mawe.201200841-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000021_mawe.201200841-Figure3-1.png",
        "caption": "Figure 3. Definitions of depth and distance of the RP part.",
        "texts": [
            " The DOE procedure consists of the following steps: (a) identify the main objective of the experiment; (b) determine the quality characteristic; (c) identify the control factors affecting the quality characteristic; (d) select an appropriate orthogonal array (OA) and assign the factor levels; (e) perform experimental trials based on the configured OA; (f) determine the optimal levels of control factors based on the S/N ratios; (g) determine the significant factors that mainly affect the quality characteristic by the analysis of variance (ANOVA); (h) verify the optimal process parameters through confirmation experiments. Four operational parameters of pH value, solvent temperature, depth of the RP, and distance of the RP part were selected as control factors to perform the Taguchi approach. The definitions of depth and distance of the RP part are shown in Figure 3. Experimental trials based on the L9 OA were performed. Support materials are removed from RP part by ultrasonic machine (frequency = 46 kHz) when the object is fabricated. Alkaline detergent (Sodium hydroxide, 1310-73-2) was used to obtain the solvent for removing support material from RP part. PH meter (pH 600) and digital thermometer (resolution = 0.1 8C) were used to monitor the variations of solvent pH value and sol- vent temperature during removing process, respectively. Three test parts were fabricated by the FDM system to evaluate the efficiency of removing process using optimal parameters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000392_j.mechmachtheory.2013.01.012-Figure3-1.png",
        "caption": "Fig. 3. Conditions of non-undercutting: a) type IIa; and b) type IIb.",
        "texts": [
            " This means that in the presence of undercutting \u2014 type IIa the tooth thickness at the bottom decreases without cutting an involute profile in the vicinity of its starting point b. When the radius \u03c13 (Fig. 2c) of the rack-cutter fillet increases considerably, the fillet gear fq crosses the radial line Ob, as well as the involute profile ba. In this case (undercutting \u2014 type IIb) besides the decrease of the tooth thickness at the bottom, the segment bq of the involute profile is also cut. The essence of undercutting \u2014 type IIa and type IIb is explained on Fig. 3, where undercutting \u2014 type I is avoided by a positive displacement of the rack-cutter at a distance Xmin. At this boundary displacement, the tip-line g-g of the rack-cutter crosses the line of action AB in starting point A. In order to define the maximum radius of the rack-cutter fillet AF, corresponding to the boundary case where there is no undercutting\u2014 type IIa, the curve \u03b7, called a boundary fillet\u2014 type IIa is drawn additionally on Fig. 3a. It is obtained as an envelope of the relative positions taken by the radial line l (the line ObE) of the gear in the plane of the rack-cutter, when realizing the meshing between the rectilinear profile AE of the rack of the involute profile ba of the gear. In other words, the profiles l and \u03b7 are also conjugated profiles at rolling without sliding of the centrode line n-n of the \u0430 rack-cutter on the reference circle of the gear of a radius r. Knowing the curve \u03b7 allows us to define the following boundary condition: the undercutting \u2014 type IIa is avoided if the real rack-cutter fillet AF is placed internally regarding the boundary fillet \u03b7 (in the material of the cutter). In Fig. 3a the rack-cutter fillet AF is placed externally regarding the curve \u03b7, as a result of which gear teeth are undercut\u2014 type IIa. Analogously the condition for non-undercutting\u2014 type IIb is defined by drawing a curve \u03be (Fig. 3b), called a boundary fillet\u2014 type IIb. In this case, the curve \u03be is obtained as a trajectory (drawn in the plane of the rack-cutter) of the point \u0430 from the plane of the reference circle of a radius r, rolling without sliding on a reference circle on the line n-n. As point b lies on the internal side of the reference circle, the trajectory drawn represents a shortened cycloid. The same curve \u03be, connected without moving with the rack-cutter, can be considered also as a conjugated curve of the starting point b of the involute profile. This means that if the real rack-cutter fillet coincides with \u03be, it will contact with point b at each moment and will not cut the involute profile bq. Therefore curve \u03be allows us to define the following boundary condition: undercutting \u2014 type IIb is avoided if the real rack-cutter fillet AF is placed internally regarding the boundary fillet \u03be. In the case shown in Fig. 3b this condition is not satisfied and as a result the gear teeth are undercut \u2014 type IIb. 3. \u0415quations of boundary fillet curves 3.1. Boundary fillet \u2014 type II\u0430 (curve \u03b7) The equations of this curve are found using the theory of plane meshing [1], where one of the two conjugated profiles is set and the other one is obtained as an envelope of the relative positions which the specified profile occupies in the plane of the searched one. In this case (Fig. 4) the specified profile is the radial line l (axis OlYl) of the gear and the searched profile is the boundary fillet \u03b7 of the rack-cutter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002380_j.ifacol.2015.08.056-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002380_j.ifacol.2015.08.056-Figure1-1.png",
        "caption": "Figure 1: Tether frame representation.",
        "texts": [
            " The xaxis points to the true North, the y-axis points the West and the z-axis points up. This frame will be denoted by the superscript n. The tether frame is a non-inertial coordinate system associated with the cardan joint mechanism. It has its origin in the point where the tether is connected with the helicopter. x and y axes rotates with respect to the fuselage of the helicopter and the z axis is always pointing towards the landing point. This frame will be denoted by the superscript t. This frame is shown in Figure 1 together with the body axes frame. Two different GNC strategies are considered in this paper for the different flight phases to be accomplished and the type of sensor information that is required to perform the mission in a safe way. The first GNC approach is the most commonly used for UAS for static platform take off and waypoint navigation in absolute coordinates. Using this approach it is possible to accomplish the mission requirements without centimeter position accuracy fusing INS/GPS information",
            " The xaxis points to the true North, the y-axis points the West and the z-axis points up. This frame will be denoted by the superscript n. Device frame: The tether frame is a non-inertial coordinate system associated with the cardan joint mechanism. It has its origin in the point where the tether is connected with the helicopter. x and y axes rotates with respect to the fuselage of the helicopter and the z axis is always pointing towards the landing point. This frame will be denoted by the superscript t. This frame is shown in Figure 1 together with the body axes frame. Figure 1: Tether frame representation. GNC approaches: Two different GNC strategies are considered in this paper for the different flight phases to be accomplished and the type of sensor information that is required to perform the mission in a safe way. The first GNC approach is the most commonly used for UAS for static platform take off and waypoint navigation in absolute coordinates. Using this approach it is possible to accomplish the mission requirements without centimeter position accuracy fusing INS/GPS information"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002965_(asce)as.1943-5525.0000730-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002965_(asce)as.1943-5525.0000730-Figure4-1.png",
        "caption": "Fig. 4. Great Planes PT-60 RC plane and its main parameters in the X-plane flight simulator",
        "texts": [
            " F or p er so na l u se o nl y; a ll ri gh ts r es er ve d. flight procedure. Some control modes and commands, such as velocity, altitude, and other path information, are sent to the autopilot directly. The autopilot, developed by the authors\u2019 work group, uses the position, orientation, heading angle, and other information of the UAV to calculate the low-level control signals. The Great Planes PT-60 RC plane was used in the simulations. The frame of the plane with its main parameters is shown in Fig. 4. The proposed method is compared with the cascaded PID controller in an autopilot, which consists of two loops: the yaw-to-roll loop and the yaw-to-rudder loop. The inputs are the aileron and the rudder, and the output is the measured yaw angle. The major character is robust to disturbances, and two wind conditions are introduced: gust wind and turbulent wind. One of the things that must be considered is the desired profile of the heading angle. There are two different cases of the heading angle: a fixed value of the angle rate and a changeable value of the angle rate"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003256_j.eng.2017.04.007-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003256_j.eng.2017.04.007-Figure9-1.png",
        "caption": "Fig. 9. Schematic model of the tri-subsystem receptance coupling via the screw-nut. Block \u201c\u00d7 d\u201d stands for multiplication by the diameter of the ball screw.",
        "texts": [
            " However, it is common for a linear scale attached at the base of the sliding component close to the working point (Coordinate 3) to be used for position feedback, so as to eliminate the negative effects of the clearance, flexible deflection, and thermal and pitch which consists of the rotational motion and vibration of the assembly of the ball-screw, the coupling, and the motor rotor with its rotational supports; subsystem B, which consists of the axial motion and vibration of the ball-screw with its axial supports; and subsystem C, which consists of the axial motion and vibration of the sliding component due to the carriage\u2019s axial and pitching movement. Fig. 9 shows the schematic model of the proposed tri-subsystem receptance coupling via the screw-nut for a ball-screw feed drive. This model was used to connect (i.e., assemble) the three subsystems mentioned above. Eq. (1) is the receptance coupling equation according to the subsystem linking configuration shown in Fig. 9, which connected the three subsystems via the screw-nut in order to acquire a coupled system: (1) a2a2 a2c2 a2c3 a2a2 c2a2 c2c2 c2c3 c2c2 c2c3 c3a2 c3c2 c3c3 c3c2 c3c3 a2a1 b a a1a1 c2c1 b1b1 c1 b 2 a c3c1 ' ' ' 0 0 ' ' ' 0 ' ' ' 0 ( ) h h h h h h h h h h h h h h h i h h h h i h = \u2212 \u2212 + + T a1a2 b a 1 c1 c1c2ja ja c3c1 1 i h i hk c h \u03c9 \u2212 \u2212 + + where i and \u03c9 are the imaginary unit and angular frequency, respectively; ha1a1, hb1b1, hc1c2 \u2026 are the pre-coupling FRFs (receptances) of the subsystems; h'a2a2, h'a2c2 \u2026 are the post-coupling FRFs of the connected system; ia b is the transmission ratio from the interface coordinate of subsystem A to that of subsystem B; and kja and cja are the equivalent axial stiffness and damping of the screw-nut joint surface, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003507_978-981-13-0305-0_5-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003507_978-981-13-0305-0_5-Figure8-1.png",
        "caption": "Fig. 8 Parameters related to cost",
        "texts": [],
        "surrounding_texts": [
            "The present stage describes the procedure for chemical post processing of the specimen. A chemical bath is prepared with acetone (Acetone 85% and distilled water 15%) as ABS plastic is soluble in acetone and also due to its low cost, low toxicity and good diffusion property. The specimen is now immersed for one hour. After removing it from ethylene, it is dried and weighed and its dimensions are measured. The surface roughness of the part is again measured to find out any difference in roughness."
        ]
    },
    {
        "image_filename": "designv11_13_0002811_mmar.2016.7575317-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002811_mmar.2016.7575317-Figure1-1.png",
        "caption": "Fig. 1. Graph of \ud835\udc39 .",
        "texts": [],
        "surrounding_texts": [
            "Let \ud835\udd4b be an isolated time scale such that sup\ud835\udd4b = \u221e and \u210e = sup{\ud835\udf07(\ud835\udc61) : \ud835\udc61 \u2208 \ud835\udd4b} < \u221e. Then we consider the following Cucker-Smale model on \ud835\udd4b:{ \ud835\udc65(\ud835\udf0e(\ud835\udc61)) = \ud835\udc65(\ud835\udc61) + \ud835\udf07(\ud835\udc61)\ud835\udc39 (\ud835\udc65(\ud835\udc61), \ud835\udc63(\ud835\udc61)) \ud835\udc63(\ud835\udf0e(\ud835\udc61)) = \ud835\udc3a(\ud835\udc65(\ud835\udc61)) \u22c5 \ud835\udc63(\ud835\udc61) (2) with initial conditions \ud835\udc63(\ud835\udc610) = ( \ud835\udc6301 , . . . , \ud835\udc63 0 \ud835\udc41 ) \u2208 \ud835\udd3c \ud835\udc41 and \ud835\udc65(\ud835\udc610) = ( \ud835\udc650 1, . . . , \ud835\udc65 0 \ud835\udc41 ) \u2208 \ud835\udd3c \ud835\udc41 , where \ud835\udc38 = \u211d 3. Observe that for a particular case, when \ud835\udd4b = \u210e\u21150, we get system (1) analyzed in [7], [8]. Definition 2. We say that system (2) tends to consensus if the consensus parameters tend to a common value, namely lim\ud835\udc5b\u2192\u221e \ud835\udc63\ud835\udc56(\ud835\udf0e \ud835\udc5b(\ud835\udc610)) = \ud835\udc63, for every \ud835\udc56 = 1, ..., \ud835\udc41 . In other words: a consensus is a state in which all agents have the same consensus parameter. Let us now consider the diagonal of \ud835\udd3c\ud835\udc41 \ud835\udcac\ud835\udc51 = {(\ud835\udc5e1, \ud835\udc5e2, ..., \ud835\udc5e\ud835\udc41 ) \u2208 \ud835\udd3c \ud835\udc41 : \ud835\udc5e1 = \ud835\udc5e2 = \u22c5 \u22c5 \u22c5 = \ud835\udc5e\ud835\udc41} and its orthogonal complement with respect to the standard inner product \ud835\udcac\u22a5 \ud835\udc51 = {(\ud835\udc5e1, \ud835\udc5e2, ..., \ud835\udc5e\ud835\udc41 ) \u2208 \ud835\udd3c \ud835\udc41 : \ud835\udc5e1 + \ud835\udc5e2 + \u22c5 \u22c5 \u22c5+ \ud835\udc5e\ud835\udc41 = 0} . Therefore, every \ud835\udc5e \u2208 \ud835\udd3c \ud835\udc41 can be uniquely written as \ud835\udc5e = \ud835\udc5e\ud835\udc51 + \ud835\udc5e\u22a5, with \ud835\udc5e\ud835\udc51 \u2208 \ud835\udcac\ud835\udc51 and \ud835\udc5e\u22a5 \u2208 \ud835\udcac\u22a5 \ud835\udc51 . Observe that convergence of \ud835\udc5e\ud835\udc56 to a common value (see Definition 2) means convergence to the diagonal of \ud835\udd3c \ud835\udc41 or, if we set \ud835\udc44 = \ud835\udd3c \ud835\udc41/\ud835\udcac\ud835\udc51, convergence to 0 in this quotient space. In \ud835\udc44 \u2243 \ud835\udcac\u22a5 \ud835\udc51 we fix an inner product \u27e8\ud835\udc5d, \ud835\udc5e\u27e9\ud835\udc44 = 1 2 \ud835\udc41\u2211 \ud835\udc56,\ud835\udc57=1 \u27e8\ud835\udc5d\ud835\udc56 \u2212 \ud835\udc5d\ud835\udc57 , \ud835\udc5e\ud835\udc56 \u2212 \ud835\udc5e\ud835\udc57\u27e9 and denote by \u2225 \u22c5 \u2225\ud835\udc44 the norm induced by this inner product. Remark 3. [14] The following definitions of consensus are equivalent (i) lim\ud835\udc5b\u2192\u221e \ud835\udc63\ud835\udc56(\ud835\udf0e \ud835\udc5b(\ud835\udc610)) = \ud835\udc63, for \ud835\udc56 = 1, ..., \ud835\udc41 ; (ii) lim\ud835\udc5b\u2192\u221e \ud835\udc63\u22a5\ud835\udc56(\ud835\udf0e \ud835\udc5b(\ud835\udc610)) = 0, for \ud835\udc56 = 1, ..., \ud835\udc41 . Let \ud835\udc4b := \ud835\udd3c \ud835\udc41/\ud835\udcac\ud835\udc51 \u2243 \ud835\udcac\u22a5 \ud835\udc51 and \ud835\udc49 := \ud835\udd3c \ud835\udc41/\ud835\udcac\ud835\udc51 \u2243 \ud835\udcac\u22a5 \ud835\udc51 . We assume a map \ud835\udc3a : \ud835\udc4b \u2192 \ud835\udc38\ud835\udc5b\ud835\udc51(\ud835\udc49 ), where \ud835\udc4b,\ud835\udc49 are endowed with a norm, what means that the expression \u2223\u2223\ud835\udc3a(\ud835\udc65)\u2223\u2223\ud835\udc44 refers to the operator norm of \ud835\udc3a(\ud835\udc65) with respect to the norm on \ud835\udc44 = \ud835\udd3c \ud835\udc41/\ud835\udcac\ud835\udc51. We also assume that \ud835\udc3a satisfies, for some \ud835\udc54 > 0 and \ud835\udefd \u2265 0, that for all \ud835\udc65 \u2208 \ud835\udc4b , \u2223\u2223\ud835\udc3a(\ud835\udc65)\u2223\u2223\ud835\udc44 \u2264 1\u2212 \u210e\ud835\udc54 (1 + \u2223\u2223\ud835\udc65\u2223\u22232\ud835\udc44)\ud835\udefd , (3) for \u210e = sup{\ud835\udf07(\ud835\udc61) : \ud835\udc61 \u2208 \ud835\udd4b} and 0 < \u210e < 1 \ud835\udc54 . We next assume \ud835\udc39 : \ud835\udc4b \u00d7 \ud835\udc49 \u2192 \ud835\udc49 to be a function (Lipschitz or \ud835\udc361) satisfying, for some \ud835\udc36, \ud835\udeff > 0 and 0 \u2264 \ud835\udefe < 1, that for \ud835\udc65 \u2208 \ud835\udc4b and \ud835\udc63 \u2208 \ud835\udc49 , \u2223\u2223\ud835\udc39 (\ud835\udc65, \ud835\udc63)\u2223\u2223\ud835\udc44 \u2264 \ud835\udc36(1 + \u2223\u2223\ud835\udc65\u2223\u22232\ud835\udc44) \ud835\udefe 2 \u2223\u2223\ud835\udc63\u2223\u2223\ud835\udeff\ud835\udc44 . (4) For simplicity of notation let us put \u039b(\ud835\udc63) := \u2223\u2223\ud835\udc63\u2223\u22232\ud835\udc44 and \u0393(\ud835\udc65) := \u2223\u2223\ud835\udc65\u2223\u22232\ud835\udc44. In the case \ud835\udc63 = \ud835\udc65 = 0 we will write shortly \u039b0 and \u03930 instead of \u039b(0) ,\u0393(0), respectively. Now we can formulate the following result. Theorem 4. Assume that \ud835\udc39 satisfies condition (4), \ud835\udc3a satisfies condition (3) and 0 < \u210e < 1 \ud835\udc54 . Assume also that one of the three following hypothesis is fulfilled: (i) \ud835\udefd < 1\u2212\ud835\udefe 2 , (ii) \ud835\udefd = 1\u2212\ud835\udefe 2 and \u039b0 < ( \ud835\udc54 \ud835\udc36\ud835\udc45(\ud835\udeff) ) 2 \ud835\udeff , (iii) \ud835\udefd > 1\u2212\ud835\udefe 2 , ( 1 a ) 1 \ud835\udefc\u22121 [( 1 \ud835\udefc ) 1 \ud835\udefc\u22121 \u2212 ( 1 \ud835\udefc ) \ud835\udefc \ud835\udefc\u22121 ] > b (5) and \u210e < [( 1\u2212 1 \ud835\udefc ) \u2212 ( 1 a\ud835\udefc )\u2212 2 \ud835\udefc\u22121 b ] \u22c5 ( \ud835\udc45(\ud835\udeff) 2\ud835\udc45(\ud835\udeff) + a ) \ud835\udc45(\ud835\udeff) a\ud835\udc54 . (6) In the above, \ud835\udefc = 2\ud835\udefd+\ud835\udefe, \ud835\udc45(\ud835\udeff) = max{1, 1 \ud835\udeff } and a = \u039b \ud835\udeff 2 0 \ud835\udc36\ud835\udc45(\ud835\udeff) \ud835\udc54 and b = 1 + \u0393 1 2 0 . Then there exists a constant \ud835\udc350 (independent of \ud835\udc61 = \ud835\udf0e\ud835\udc5b(\ud835\udc610) and given explicitly in the proof of each of the three cases), such that\u2223\u2223\u2223\u2223\ud835\udc65(\ud835\udf0e\ud835\udc5b(\ud835\udc610)) \u2223\u2223\u2223\u22232 \ud835\udc44 \u2264 \ud835\udc350 for all \ud835\udc5b \u2208 \u2115. Additionally,\u2223\u2223\u2223\u2223\ud835\udc63(\ud835\udf0e\ud835\udc5b(\ud835\udc610)) \u2223\u2223\u2223\u2223 \ud835\udc44 \u2192 0, when \ud835\udc5b \u2192 \u221e. Moreover, there exists ?\u0302? \u2208 \ud835\udc4b such that \ud835\udc65\u22a5(\ud835\udf0e\ud835\udc5b(\ud835\udc610)) \u2192 ?\u0302?, when \ud835\udc5b \u2192 \u221e. Proof: The proof is similar to the one given by Cucker and Smale in [8]. Throughout the proof we will write \u0393(\ud835\udc5b) instead of \u0393(\ud835\udc65(\ud835\udf0e\ud835\udc5b(\ud835\udc610))) = \u2223\u2223\ud835\udc65(\ud835\udf0e\ud835\udc5b(\ud835\udc610))\u2223\u22232. Let us set initial conditions \ud835\udc65(\ud835\udc610) = \ud835\udc650 and \ud835\udc63(\ud835\udc610) = \ud835\udc630. Then we observe that \u2223\u2223\ud835\udc63(\ud835\udf0e\ud835\udc5b(\ud835\udc610))\u2223\u2223\ud835\udeff\ud835\udc44 = \u2223\u2223\ud835\udc3a(\ud835\udf0e\ud835\udc5b\u22121(\ud835\udc610))\ud835\udc63(\ud835\udf0e \ud835\udc5b\u22121(\ud835\udc610))\u2223\u2223\ud835\udeff\ud835\udc44 \u2264 = \u2223\u2223\ud835\udc63(\ud835\udc610)\u2223\u2223\ud835\udeff\ud835\udc44 \ud835\udc5b\u22121\u220f \ud835\udc56=0 ( 1\u2212 \ud835\udc54\u210e (1 + \u0393(\ud835\udc56))\ud835\udefd )\ud835\udeff . (7) Now let \ud835\udc5b\u2217 = max{\u2223\u2223\ud835\udc65(\ud835\udf0e\ud835\udc56(\ud835\udc610))\u2223\u2223\ud835\udc44 : \ud835\udc56 = 0, 1, ..., \ud835\udc5b}, so it is a point that maximizes \u0393(\ud835\udc56). Then for all \ud835\udc58 \u2264 \ud835\udc5b we evaluate: \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc58(\ud835\udc610)) \u2225\ud835\udc44\u2264\u2225 \ud835\udc65(\ud835\udc610) \u2225\ud835\udc44 + \ud835\udc58\u22121\u2211 \ud835\udc56=0 \u210e\ud835\udc36 ( 1+ \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc57(\ud835\udc610)) \u22252\ud835\udc44 ) \ud835\udefe 2 \u2225 \ud835\udc63(\ud835\udf0e\ud835\udc57(\ud835\udc610)) \u2225\ud835\udeff\ud835\udc44 \u22c5 \ud835\udc57\u22121\u220f \ud835\udc56=0 ( 1\u2212 \ud835\udc54\u210e (1 + \u0393(\ud835\udc56))\ud835\udefd )\ud835\udeff \u2264\u2225 \ud835\udc65(\ud835\udc610) \u2225\ud835\udc44 + \ud835\udc36 \u2225 \ud835\udc63(\ud835\udc610) \u2225\ud835\udeff\ud835\udc44 (1 + \u0393(\ud835\udc5b\u2217)) \ud835\udefe 2 \u22c5 (1 + \u0393(\ud835\udc5b\u2217))\ud835\udefd \ud835\udc54 \ud835\udc45(\ud835\udeff) , where \u210e = sup{\ud835\udf07(\ud835\udc61), \ud835\udc61 \u2208 \ud835\udd4b} and \ud835\udc45(\ud835\udeff) = max{1, 1 \ud835\udeff }. Putting \ud835\udc58 = \ud835\udc5b\u2217 we get (1 + \u0393(\ud835\udc5b\u2217)) 1 2 \u2264 ( 1 + \u0393 1 2 0 ) + \ud835\udc36\u039b \ud835\udeff 2 0 \ud835\udc45(\ud835\udeff) \ud835\udc54 (1 + \u0393(\ud835\udc5b\u2217))\ud835\udefd+ \ud835\udefe 2 . (8) Now we will introduce an auxiliary function \ud835\udc39 (\ud835\udc67) := \ud835\udc67\u2212a\ud835\udc672\ud835\udefd+\ud835\udefe\u2212b, where \ud835\udc67 = (1 + \u0393(\ud835\udc5b\u2217)) 1 2 , a = \ud835\udc36\u039b \ud835\udeff 2 0 \ud835\udc45(\ud835\udeff) \ud835\udc54 and b = 1 + \u0393 1 2 0 . By (8) we have \ud835\udc39 (\ud835\udc67) \u2264 0. (i) Assume \ud835\udefd < 1\u2212\ud835\udefe 2 . Since \ud835\udc39 (\ud835\udc67) \u2264 0 implies (1 + \u0393(\ud835\udc5b\u2217)) 1 2 \u2264 \ud835\udc480 with \ud835\udc480 = max{(2a) 1\u2212\ud835\udefe 1\u2212\ud835\udefe\u22122\ud835\udefd , 2b} (see Lemma 1 in [7]) and \ud835\udc480 is independent of \ud835\udc5b, it follows that, for all \ud835\udc5b \u2265 0, \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc5b\u2217 (\ud835\udc610)) \u22252\ud835\udc44\u2264 \ud835\udc350 = \ud835\udc482 0 \u2212 1. Therefore, using (7), we get that expression \u2223\u2223\ud835\udc63(\ud835\udf0e\ud835\udc5b(\ud835\udc610))\u2223\u2223\ud835\udc44 \u2264 \u2223\u2223\ud835\udc63((\ud835\udc610)\u2223\u2223\ud835\udc44 ( 1\u2212 \ud835\udc54\u210e (1+\ud835\udc350)\ud835\udefd )\ud835\udc5b tends to zero when \ud835\udc5b \u2192 \u221e. (ii) Assume \ud835\udefd = 1\u2212\ud835\udefe 2 . Then inequality \ud835\udc39 (\ud835\udc67) \u2264 0 takes the form \ud835\udc67 ( 1\u2212 \ud835\udc36\u039b \ud835\udeff 2 0 \ud835\udc45(\ud835\udeff) \ud835\udc54 ) \u2264 1 + \u0393 1 2 0 . By hypothesis expression between parenthesis is positive, thus \ud835\udc67 \u2264 \ud835\udc480 = 1+\u0393 1 2 0\u239b \u239d1\u2212\ud835\udc36\u039b \ud835\udeff 2 0 \ud835\udc45(\ud835\udeff) \ud835\udc54 \u239e \u23a0 and we proceed as in case (i). (iii) Assume \ud835\udefd > 1\u2212\ud835\udefe 2 . Let \ud835\udefc = 2\ud835\udefd + \ud835\udefe, then since \ud835\udc39 (\ud835\udc67) = \ud835\udc67 \u2212 a\ud835\udc672\ud835\udefd+\ud835\udefe \u2212 b, we get \ud835\udc39 \u2032(\ud835\udc67) = 1\u2212\ud835\udefca\ud835\udc67\ud835\udefc\u22121. The standard calculus shows that \ud835\udc39 \u2032(\ud835\udc67) = 0 if and only if \ud835\udc67\u2217 = ( 1 \ud835\udefca ) 1 \ud835\udefc\u22121 and that \ud835\udc39 (\ud835\udc67\u2217) = ( 1 a ) 1 \ud835\udefc\u22121 [( 1 \ud835\udefc ) 1 \ud835\udefc\u22121 \u2212 ( 1 \ud835\udefc ) \ud835\udefc \ud835\udefc\u22121 ] \u2212 b. By assumption (5) we have \ud835\udc39 (\ud835\udc67\u2217) > 0. Since \ud835\udc39 (0) = \u2212b, the graph of \ud835\udc39 is the following For \ud835\udc5b \u2208 \u2115 let \ud835\udc67(\ud835\udc5b) = (1 + \u0393(\ud835\udc5b\u2217)) 1 2 , where \ud835\udc5b\u2217 = max{\u2225 \ud835\udc65(\ud835\udf0e\ud835\udc56(\ud835\udc610)) \u2225\ud835\udc44: \ud835\udc56 = 0, 1, ..., \ud835\udc5b}. If \ud835\udc5b = 0, then \ud835\udc5b\u2217 = 0 and \ud835\udc67(0) = (1 + \u03930) 1 2 \u2264 1 + \u0393 1 2 0 = b < ( 1 a ) 1 \ud835\udefc\u22121 ( 1 \ud835\udefc ) 1 \ud835\udefc\u22121 = \ud835\udc67\u2217. And since \ud835\udc39 (\ud835\udc67\ud835\udc59) = 0 and \ud835\udc39 (\ud835\udc67(0)) = \ud835\udc67(0) \u2212 a\ud835\udc67(0)\ud835\udefc \u2212 b, we get that \ud835\udc39 (\ud835\udc67(0)) \u2264 \ud835\udc39 (\ud835\udc67\ud835\udc59) implies \ud835\udc67(0) \u2264 \ud835\udc67\ud835\udc59. Assume now that there exists \ud835\udc5b \u2208 \u2115 such that \ud835\udc67(\ud835\udc5b) \u2265 \ud835\udc67\ud835\udc62 and let \ud835\udc41 be the first such \ud835\udc5b. Then \ud835\udc41 = \ud835\udc41\u2217 \u2265 1 and for all \ud835\udc5b < \ud835\udc41 it holds (1+\u0393(\ud835\udc5b)) 1 2 \u2264 \ud835\udc67(\ud835\udc41 \u2212 1) \u2264 \ud835\udc67\ud835\udc59. Thus for \ud835\udc5b < \ud835\udc41 we have \u0393(\ud835\udc5b) \u2264 \ud835\udc350 = \ud835\udc672\ud835\udc59 \u2212 1, in particular, \u0393(\ud835\udc41 \u2212 1) \u2264 \ud835\udc350. What more, since \ud835\udc67(\ud835\udc41) \u2265 \ud835\udc67\ud835\udc59 we get \u0393(\ud835\udc41) \u2265 \ud835\udc672\ud835\udc62 \u2212 1. The latter inequality implies that \u0393(\ud835\udc41)\u2212\u0393(\ud835\udc41\u22121) \u2265 \ud835\udc672\ud835\udc62\u2212\ud835\udc672\ud835\udc59 \u2265 \ud835\udc672\u2217\u2212\ud835\udc672\ud835\udc59 \u2265 (\ud835\udc67\u2217\u2212\ud835\udc67\ud835\udc59)\ud835\udc67\u2217 . (9) From the intermediate value theorem, there exists \ud835\udc50 \u2208 [\ud835\udc67\ud835\udc59, \ud835\udc67\u2217] such that \ud835\udc39 \u2032(\ud835\udc50) = \ud835\udc39 (\ud835\udc67\u2217)\u2212\ud835\udc39 (\ud835\udc67\ud835\udc59) \ud835\udc67\u2217\u2212\ud835\udc67\ud835\udc59 . And since \ud835\udc39 \u2032(\ud835\udc50) \u2265 0 and \ud835\udc39 \u2032(\ud835\udc50) = 1\u2212 a\ud835\udefc\ud835\udc50\ud835\udefc\u22121, we get \ud835\udc67\u2217 \u2212 \ud835\udc67\ud835\udc59 \u2265 \ud835\udc39 (\ud835\udc67\u2217). Using (9) we write \u0393(\ud835\udc41)\u2212 \u0393(\ud835\udc41 \u2212 1) \u2265 \ud835\udc67\u2217\ud835\udc39 (\ud835\udc67\u2217) . (10) But \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc41 (\ud835\udc610)) \u2225\ud835\udc44 \u2212 \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc41\u22121(\ud835\udc610)) \u2225\ud835\udc44\u2264 \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc41 (\ud835\udc610))\u2212 \u2225\ud835\udc44 \u2212 \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc41\u22121(\ud835\udc610)) \u2225\ud835\udc44 \u210e\ud835\udc36 ( 1+ \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc41\u22121(\ud835\udc610)) \u22252\ud835\udc44 ) \ud835\udefe 2 \u2225 \ud835\udc63(\ud835\udf0e\ud835\udc41\u22121(\ud835\udc610)) \u2225\ud835\udeff\ud835\udc44 \u210e\ud835\udc36 ( 1+ \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc41\u22121(\ud835\udc610)) \u22252\ud835\udc44 ) \ud835\udefe 2 \u2225 \ud835\udc63((\ud835\udc610)) \u2225\ud835\udeff\ud835\udc44 \u22c5 \ud835\udc41\u22121\u220f \ud835\udc56=0 ( 1\u2212 \ud835\udc54\u210e (1 + \u0393(\ud835\udc56))\ud835\udefd )\ud835\udeff \u2264 \u210e\ud835\udc36 (1 +\ud835\udc350) \ud835\udefe 2 \u039b \ud835\udeff 2 0 . Therefore, \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc41 (\ud835\udc610)) \u22252\ud835\udc44 \u2212 \u2225 \ud835\udc65(\ud835\udf0e\ud835\udc41\u22121(\ud835\udc610)) \u22252\ud835\udc44\u2264( 2\ud835\udc35 1 2 0 + \u210e\ud835\udc36(1 +\ud835\udc350) \ud835\udefe 2 ) \u210e\ud835\udc36(1 +\ud835\udc350) \ud835\udefe 2 \u039b \ud835\udeff 2 0 , which together with (10) gives \ud835\udc67\u2217\ud835\udc39 (\ud835\udc67\u2217) \u2264 ( 2\ud835\udc35 1 2 0 + \u210e\ud835\udc36(1 +\ud835\udc350) \ud835\udefe 2 ) \u210e\ud835\udc36(1 +\ud835\udc350) \ud835\udefe 2 \u039b \ud835\udeff 2 0 . Now using definition of \ud835\udc39 (\ud835\udc67\u2217), inequality \ud835\udc350 \u2264 \ud835\udc672\u2217 \u2212 1 and since \ud835\udc67\u2217 = ( 1 \ud835\udefca ) 1 \ud835\udefc\u22121 we write, ( 1 a ) 2 \ud835\udefc\u22121 [( 1 \ud835\udefc ) 2 \ud835\udefc\u22121 \u2212 ( 1 \ud835\udefc )\ud835\udefc+1 \ud835\udefc\u22121 ] \u2212 b \u2264 ( 2 + \ud835\udc36\u039b \ud835\udeff 2 0 \ud835\udc54 ) \u210e\ud835\udc36\u039b \ud835\udeff 2 0 ( 1 \ud835\udefca ) 2 \ud835\udefc\u22121 , which together with a = \ud835\udc36\u039b \ud835\udeff 2 0 \ud835\udc45(\ud835\udeff) \ud835\udc54 implies that we can rewrite the last inequality in the following equivalent form( 1\u2212 1 \ud835\udefc ) \u2212 ( 1 a\ud835\udefc )\u2212 2 \ud835\udefc\u22121 b \u2264( 2 + \ud835\udc4e \ud835\udc45(\ud835\udeff) ) \u210ea\ud835\udc54 \ud835\udc45(\ud835\udeff) , which contradicts hypothesis (6). Therefore, we can conclude that for all \ud835\udc5b \u2208 \u2115, \ud835\udc67(\ud835\udc5b) \u2264 \ud835\udc67\ud835\udc59 and hence, \u2223\u2223\ud835\udc65(\ud835\udf0e\ud835\udc5b(\ud835\udc610))\u2223\u22232\ud835\udc44 \u2264 \ud835\udc350, which brings us to the case (i) again. And the proof is finished."
        ]
    },
    {
        "image_filename": "designv11_13_0002173_detc2014-35099-Figure21-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002173_detc2014-35099-Figure21-1.png",
        "caption": "Figure 21 Tessellated CAD model approximation of volumetric error with adaptive slicing (a, b & c)",
        "texts": [
            " The models illustrate the variation in slice thicknesses in adaptive slicing procedure of non-axisymmetric and free form models, showing how variation in slice thicknesses can occur and how the new method minimize volumetric error. In the models shown in Fig. 18, tessellated CAD model approximations of volumetric error with uniform slicing have been presented. Figures 19a, b & c shows uniformly sliced CAD models. The minimum and maximum slice thicknesses were set to 0.127mm and 0.33mm respectively and the volumetric error was set to 0.18mm 3 . Figure 20 shows graphical output related to adaptive slicing of the sample 3D models. As shown in Fig. 19 and Fig. 21, developed volumetric error based slicing methodology and corresponding program is able to slice non-axisymmetric (2.5D,extrusion) and free form models with a uniform slice thickness and as well as adaptive slicing. 9 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Current status of research in layered manufacturing demands for effective volumetric error control parameter, as it has been found that cusp height does not have any direct control on volumetric error and more important that with a very little variation in cusp height a large variation in volumetric error occurs on steep slopes of surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000209_dscc2013-3740-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000209_dscc2013-3740-Figure7-1.png",
        "caption": "Figure 7. BASE OF A DEFECTIVE CONICAL SHAPE MANIPULATOR WITH A SMALL INACCURACY (\u03b4) WITH RESPECT TO THE PLANE PASSING FROM THE MIDLINE AND CABLE T2",
        "texts": [
            " The authors of [8] explain that the tip errors depend on some inaccuracies introduced during the building phase, when cables are 7 Copyright \u00a9 2013 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/23/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use embedded in silicone. Probably after silicone polymerization the path of the embedded cables is not perfectly linear and coplanar with the midline. By means of the general mechanical model for tendon driven manipulators [22], it is possible to simulate a situation where the cables are not perfectly coplanar, as shown in Fig. 7, in which a small inaccuracy \u03b4 = 0.45mm is taken into account. In this case, the torsion \u03c4(s) and the curvature \u03be(s) are small but not equal to zero, as it occurs in a real prototype. For example, the assumed \u03b4 value generates a tip error with respect to the ideal model of 4.16%, applying the same cable tensions T1 = 6, T2 = 0 of test performed in [8], where the authors registered a relative tip error of 6.38%. In this situation, the chosen \u03b4 value takes into account a real scenario in which tip errors was partially created by the defective prototype"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001084_j.mechmachtheory.2017.05.017-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001084_j.mechmachtheory.2017.05.017-Figure6-1.png",
        "caption": "Fig. 6. A further longitudinal modification of a pinion.",
        "texts": [
            " If so, the parameter L 1 becomes a given value, and there are still two given values ( \u03d51 and L 1 ) and five unknown values ( u 1 , l 1 , \u03d52 , u 2 , l 2 ), then, Eq. (3) can be still solved. And in the optimization problem (10) , the optimization variables \u03b4\u03d52 can be replaced with L 1 for the second opti- mization, what is more, the L (i ) 1 ( i = 1 , . . . , n ) corresponding to the ideal TE can be used as the initial value in the second optimization. The longitudinal modification can be applied on either a pinion or a gear. Pinion is selected here and the design method of the longitudinal modification is described in the following. As shown in Fig. 6 , a 2D curve is designed for the further longitudinal modification. To mitigate the influence on PPTE, the further tooth modification is only applied on the regions of both ends of the tooth face width, and without any change on the concave region. In the Fig. 6 , modification curves are second degree parabola; y 1 and y 2 are the modification values; y 3 is the unmodified length in the direction of tooth face width. The principles of designing modified surfaces with a further longitudinal modification are in the following. The Non- Uniform Rational B-Splines (NURBS) is used herein. 1) The tooth surface with a further longitudinal modification is represented by a sum of two vector functions [29] , and this step will be carried after Eqs. (1 ) and ( 2 ) and before Eqs",
            " \u23a7 \u23aa \u23a8 \u23aa \u23a9 \u21c0 R ( u 1 , l 1 , L 1 , \u03b81 ) = \u21c0 r 1 ( u 1 , l 1 , L 1 , \u03b81 ) + \u03b4F ( x, y ) \u21c0 n 1 ( u 1 , l 1 , L 1 , \u03b81 ) \u21c0 N ( u 1 , l 1 , L 1 , \u03b81 ) = ( \u2202 \u21c0 r 1 \u2202 u 1 + \u2202 \u03b4F \u2202 u 1 \u21c0 n 1 + \u2202 \u21c0 n 1 \u2202 u 1 \u03b4F ) \u00d7 ( \u2202 \u21c0 r 1 \u2202 l 1 + \u2202 \u03b4F \u2202 l 1 \u21c0 n 1 + \u2202 \u21c0 n 1 \u2202 l 1 \u03b4F ) (11) here, { x = \u221a r 2 1 x + r 2 1 y y = r 1 z \u23a7 \u23aa \u23a8 \u23aa \u23a9 \u2202 \u03b4F \u2202 u 1 = \u2202 \u03b4F (x, y ) \u2202x \u2202x \u2202 u 1 + \u2202 \u03b4F (x, y ) \u2202y \u2202y \u2202 u 1 \u2202 \u03b4F \u2202 l 1 = \u2202 \u03b4F (x, y ) \u2202x \u2202x \u2202 l 1 + \u2202 \u03b4F (x, y ) \u2202y \u2202y \u2202 l 1 (1) The first vector function \u21c0 r 1 ( u 1 , l 1 , L 1 , \u03b81 ) is giv en after L (i ) 1 ( i = 1 , . . . , n ) is deriv ed; r 1 x , r 1 y , r 1 z are all the coordi- nates of \u21c0 r 1 ( u 1 , l 1 , L 1 , \u03b81 ) ; (2) A 3D grid (a network of points P ( x, y, \u03b4F ) of a further longitudinal modification) is developed, as shown in Fig. 6 ; (3) The interpolation by a bi-cubic NURBS is used, and the points P ( x, y, \u03b4F ) are fitted into a smooth surface; (4) The second vector function \u03b4( x, y ) \u21c0 n 1 ( u 1 , l 1 , L 1 , \u03b81 ) is determined by the further correction value \u03b4F ( x, y ) and the unit normal vectors \u21c0 n 1 ( u 1 , l 1 , L 1 , \u03b81 ) ; (5) Both vector functions mentioned above are represented by the same coordinates, and then the new modified tooth surface is constructed, which will be used in the TCA calculation. In order to distinguish the optimization variables L (i ) 1 ( i = 1 , "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000616_aim.2012.6265962-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000616_aim.2012.6265962-Figure13-1.png",
        "caption": "Fig. 13. Verification of visual feedback",
        "texts": [
            " 11 denotes that ZMP moves forward and reaches the edge of supporting-foot while the other foot in the air, meaning that the robot is tipping over, which does not appear in ZMP-based walking. We think that this kind of natural walking is caused by the effect of visual feedback as shown in the following subsection. We assume that two patterns of supporting-foot\u2019s contacting and input torques based on Eq. (24). Since state of Fig. 12 (a) meaning surface-contacting is thought to be a manipulator fixed at the ground as shown in Fig. 13 (a), it is clear that Eq. (24) can lift the robot\u2019s head up toward desired position. On the other hand, effectiveness of visual feedback is unclear in toe-contacting phase because there is no input to toe\u2019s joint that means the robot\u2019s non-holonomic dynamics include constraint condition of toe\u2019s joint. However, Fig. 13 (b) simulating the state of Fig. 12 (b) indicates that although one link corresponding to the foot falls by gravity, the others are pulled toward the desired position. Therefore, we can say that visual feedback may make the whole dynamics stable partially even though non-holonomic constraint be added to. Here, we discuss whether Eq. (24) makes the humanoid\u2019s pose stable. By changing the value of feedback gain Kp, the strength of force lifting the robot\u2019s head is adjustable. Here, to verify some effects that the strength of visual feedback gives to the humanoid\u2019s walking, we confirmed walkings by using some kinds of \u03b1Kp (\u03b1 is weight coefficient)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003093_s1068366617030163-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003093_s1068366617030163-Figure4-1.png",
        "caption": "Fig. 4. Pressure profiles in a cylindrical bearing (a) without misalignment, (b) with a small misalignment, and (c) with a large misalignment.",
        "texts": [
            " The multilobe design is considered in the hydrodynamic calculations by calculating and presetting the additional components of a gap. These deviations from a circular cylindrical shape refer to the designed deviations. In Fig. 3, the dimensional profiles of the additional components of gaps in two-lobe (lemon bore) and three-lobe bearings are shown. The misalignment of the shaft axis with respect to the bush axis was also modeled as a rotation of an undeformable shaft in an undeformable bush. In the absence of the misalignment, the pressure profile is symmetric with respect to a bearing (Fig. 4a); the misalignment disturbs the symmetry (Figs. 4b, 4c). A significant misalignment causes the appearance of the second f luid wedge on the opposite surface of a bush (Fig. 4c). The misalignment is modeled similarly by presetting the additional components of a gap. The performed finite-element calculation proved that the pressure profile is nonsymmetric with respect to the center of the transverse cross section of a bearing. 244 JOURNAL OF FRICTION AND WEAR Vol. 38 No. 3 2017 ZERNIN et al. These factors (multilobe bore and the misalignment of axes), along with other types of deviations from a cylindrical shape should be considered in the description of a gap between the working surfaces of a bearing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002478_s13239-016-0257-y-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002478_s13239-016-0257-y-Figure3-1.png",
        "caption": "FIGURE 3. Diagram illustrating the Spyder Catheter design.",
        "texts": [
            " This would be the catheter size required for testing in a murine model. This catheter design utilizes pressure response outlets to deliver PEG and MB evenly. There were several designs that led to prototypes. The lead design was the \u2018\u2018Spyder\u2019\u2019 catheter. Any catheter design with a \u2018\u2018cuff\u2019\u2019 around the nerve wound is not extractable without surgery because pulling the catheter out will also destroy the nerve repair. The Spyder catheter branches into multiple spindle-like tubes, forming multiple infusion regions that later converge. This shape (Fig. 3) allowed the pressure-controlled catheter to envelope the site of injury, providing an evenly distributed flow of nerve repair solution. Each spindle of this design was able to release solution in a controlled and even manner through evenly spaced slits. The design was not be more costly than other catheters types already on the market. This catheter design comes packaged with PEG and MB in a single package, so they can be opened and used together. It will be marketed as a \u2018\u2018peripheral nerve repair system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000867_physreve.89.052503-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000867_physreve.89.052503-Figure6-1.png",
        "caption": "FIG. 6. (Color online) Diagram of the experimental setup for photothermal switching.",
        "texts": [
            " Although the absolute values of both conductivities (\u03c3|| and \u03c3\u22a5) in the nanocomposites are much lower than the conductivity of pure PANI nanofibers [37], the relative enhancement of \u03c3 in the composites is large compared to that of the pristine liquid crystal. Hence these nanocomposites can be used as biconducting materials in which two conductivities can be obtained simultaneously, and switching between them is possible by an appropriate external stimulus. Here we demonstrate a light driven switching of conductivity by doping a small amount (0.5 wt %) of laser dye [4-dicyanomethylene-2-methylene-2-methyl-6-(pdimethylaminostyryl)-4H-pyran (DCM)]. A schematic diagram of the experimental setup is shown in Fig. 6. We used a tunable Ar-ion laser (130 mW) whose wavelength range falls in the absorption range of DCM. The light beam was expanded to illuminate the cell by an appropriate combination of lenses. First, the sample was cooled below 47 \u25e6C to obtain a homeotropic state (state \u03b2 in Fig. 2); then it was heated up to 53 \u25e6C, which is just about 1 \u25e6C below the homeotropic to planar anchoring transition temperature. The cell is exposed to laser light for a few seconds (2\u20134 s), and the dye molecules absorb the light energy; as a result the temperature of the exposed area is increased"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000211_1.4846195-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000211_1.4846195-Figure1-1.png",
        "caption": "FIG. 1. (Color online) Line drawings of the two parts of the clamping holder, with photo of the 3D-printed parts in titanium (separate) and silver (assembled). US penny coin (19 mm diameter) shown for scale.",
        "texts": [
            "00 VC 2014 American Vacuum Society 023201-1 Redistribution subject to AVS license or copyright; see http://scitation.aip.org/termsconditions. Download to IP: 85.220.46.231 On: Thu, 08 May 2014 03:09:25 0.1 mA afterward. The analog output from the gauge controller was recorded every 33 s using a USB-1608FS analog-digital converter from Measurement Computing. The pieces used for testing were designed as a clamping holder used to facilitate the transfer of tip carriers to and from our scanning tunneling microscope. The design used two different parts, shown in Figure 1, with a total surface area of 9.2 cm2. The first material used for construction was sterling silver (92.5% silver and 7.5% copper) and was ordered from Shapeways (www.shapeways.com). This ultimately produces a cast part using a \u201clost wax\u201d process, where high-detail printing is done in wax, a negative mold created from the master, and metal casting done after the wax is melted or burned out. The other material, titanium, was printed using direct laser heating to sinter metal powder into a solid part; this was ordered from i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003348_1350650118758742-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003348_1350650118758742-Figure2-1.png",
        "caption": "Figure 2. Normal and tangential contact pressure for (a) elastically deformed asperity; (b) plastically deformed asperity.",
        "texts": [
            " When the tangential loads are not enough to cause gross sliding, the response of each asperity is coupled to the normal loads, i.e. the shorter the asperity height, the less normal loads the asperity carries and hence that slides more easily compared to a taller one. The Mindlin partial slip solution31 could be used to calculate the tangential contact loads prior to the inception of gross sliding under the condition of elastic deformation. The contact area can be divided into stick and slip regimes (shown in Figure 2(a)). The formula characterizing the relationship between tangential contact loads and displacement can be expressed as26 Qe \u00bc fWe 1 1 4G fE! 3=2 !5!e 1 04!5!e 8>< >: \u00f011\u00de where is the tangential displacement, !e is the normal interference for the stick-slip inception of elastically deformed asperities, !e \u00bc 4G fE , and G is the combined shear modulus. Furthermore, apart from the Mindlin theories, the Ciavarella\u2013Jager theorem32,33 also can be used to solve the partial slip as a superposition of two normal contact problems. For asperities under plastic deformation, the pressure over the contact area is uniform with a constant KH. The contact regime of these asperities has only two states, either slip or stick. Fujimoto26 gave the relationship between the tangential force and displacement combining experimental test and theoretical analysis (shown in Figure 2(b)), i.e. Qp \u00bc 2 GR1=2 2! !c\u00f0 \u00de 1=2 !5!p fRKH 2! !c\u00f0 \u00de !5!p ( \u00f012\u00de where !p is the normal interference for the stickslip inception of plastically deformed asperities, !p \u00bc 2G2 2 f2K2H2R \u00fe !c 2 . A statistical summation method was used to compute the total tangential loads. The tangential contact loads for rough surfaces may be normalized by dividing by fAnE, Furthermore, all the length parameters and variables are normalized by . It should be noted that equation (11) is applicable only under elastic contact conditions, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001985_ecce.2015.7309702-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001985_ecce.2015.7309702-Figure6-1.png",
        "caption": "Fig. 6. Motor B: (a) precision-machined drive/non-drive end journal bearings with 75 \u03bcm and 90 \u03bcm clearance; (b) design drawing of adapter and",
        "texts": [
            " It is very important to ensure that the oil is supplied evenly throughout the bearing-shaft contact for maintaining the oil film for stable operation. To direct the oil to the center of the bearing axial length, an adapter was added to the end shield to guide the oil to the top of the journal bearing and shaft contact, as shown in Figs. 5 and 6(b). Holes were drilled in the axial and radial direction in the upper part of the bearings to guide the incoming oil from the adapter to the shaft-bearing clearance, as shown in Fig. 6(a). The path of the oil flow through the inlet, channel, adapter, bearing, shaft, channel, and outlet is shown in Fig 5 and in the bearing design drawing shown in Fig. 6(b). It is also critical to control the \u201cflow\u201d of oil to maintain the oil film in the bearing for operation of the motor within its thermal limits. The flow of ISO viscosity grade 68 oil was controlled by feeding the oil into the grease inlet of the end shield through a valve, as shown in Fig. 7, and the oil leaving the grease outlet was circulated back to the inlet. Rubber seals were placed on the shaft to minimize the leakage of oil. The motor was fixed to a 42.0 kg steel baseplate to eliminate other sources of vibration due to structural looseness of the foundation, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003705_ecce.2018.8558256-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003705_ecce.2018.8558256-Figure6-1.png",
        "caption": "Fig. 6. In the second simulation of the proposed procedure, the rotor q\u2212axis current is: irq = \u2212isq(M/Lr). The rotor flux space vector lies along the d\u2212axis, the RFO model equation can be used to get the motor torque and the rotor slip angular frequency.",
        "texts": [
            " Using the motor inductances, computed in the first step of the procedure, the condition (11) can be imposed. In the second FEA of the procedure, stator and rotor current vectors are: is = isd + jisq ir,2 = \u2212jisq M Lr (22) where the stator current is the same as the first step (18), whilst the rotor current has to be corrected. The rotor induced current can be considered uniformly distributed within the rotor bars only when the motor is working at low slip frequency, e.g. the normal operating condition, fed by the grid. In the example reported in Fig. 6, the FE field solution is shown when currents vector in (22) are imposed in the stator and rotor windings. The currents isd and isq are the same as in the the first step. In the second FEA of the procedure, the relationship (11) is applied. Looking at RFO field solution in Fig. 6, the rotor flux lines are almost parallel to the d-axis, and the difference with the field map in Fig. 5 can be observed, in which the rotor flux lines widely cross the d-axis. Thus, the rotor q-axis flux has been significantly reduced (in that example the torque current is three times the rated one). In Table I FEA results are reported in several working points, increasing the motor torque current. In each of them, using the proposed strategy, \u03bbrq,2, in the second simulation is almost equal to zero"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003148_j.triboint.2017.08.010-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003148_j.triboint.2017.08.010-Figure5-1.png",
        "caption": "Fig. 5. Film shapes calculated with HS and QS models and interferogram excerpted from Refs. [2,3] (U* \u00bc 82 10 11).",
        "texts": [
            " The selected case for the present comparison is an unblended conical roller (cone angle: 7.9 ) on a glass disc from Refs. [2,3]. To simplify the calculation, the case was reduced into the model of a cylindrical roller of the mean radius in a lubricated rolling contact with a semi-infinite plane (the contact type of QS/HS, as shown in Fig. 4(a)). To show the free edge surface effect, the deformation of the roller was calculated with both QS and HS models. The calculated EHL results were compared with the experimental data. The results are shown in Figs. 5 and 6. Fig. 5(a) shows the interferogram with U* \u00bc 82 10 11 extracted from Refs. [2,3]. Fig. 5(b) and (c) show the film shapes calculated using the HS and QS models. Both calculated film shapes are consistent with the experiment data (Fig. 5(a)). Fig. 6 shows the comparison of the measured film thickness [2,3] and the calculated film thickness with the two different models along the end closure in the axial direction and the centreline. Differences in the calculated film thickness between the two models are considerably small and they are consistent with the measured data of Wymer and Cameron [2,3]. That the two models produce almost the same film thickness on the centreline is because it is far from the free end surface. However, the fact that the film shape (Fig. 5) and thickness (Fig. 6) near the end surface are also almost identical using the two models remains interesting. The free end surface does not seem to affect the behaviour of finite line contact EHL. However, it should be noticed that the contact was between a glass disc and a steel roller in Wymer and Cameron's experiment [2,3]. Moreover, the main deformation occurred on the glass disc, which has only approximately one-third of the Young's modulus of steel. The glass disc can be considered as a semi-infinite body and its deformation can be calculated using the HS model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003380_s11665-018-3162-8-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003380_s11665-018-3162-8-Figure9-1.png",
        "caption": "Fig. 9 Wear mechanism of biomimetic sample: (a) the first stage of wear (b) the second stage of wear",
        "texts": [
            " Additionally, among the laser-treated samples, the weight loss of the samples demonstrated a decreasing tendency followed by an increase with increasing laser energy input. In other words, when the laser energy density was 420.1 J/cm2, No. 3 sample had the lowest weight loss of 28 mg, as one-third of the untreated sample (85 mg), along with the most effective wear resistance. According to previous studies (Ref 22, 27), the reciprocating line wear of the 7075 aluminum alloy has specific characteristics. As presented in Fig. 9, the wear process was divided into two stages. Figure 9(a) shows that at the beginning of the abrasion, the substrate and the bionic units were in the same plane, similar to the untreated sample. Following, in the second-stage abrasion, the softer substrate presented abrasion first, whereas the harder unit body was less worn. As shown in Fig. 9(b), the unit played the role of a dam that inhibited wear in the second stage. Table 2 and Fig. 7 show that the unit hardness of No. 3 was maximal, up to 165.7+12 9HV and 56% higher than the substrate hardness (106.3+5 -6HV), whereas the grain size of sample No. 3 was the finest and most uniform. Therefore, the role of the dam played in sample No. 3 was the best, proving that it had good wear resistance. Also, compared to samples No. 1 and No. 5, No. 5 had a lower hardness and larger grain size, but lower weight loss"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001808_1.4037667-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001808_1.4037667-Figure3-1.png",
        "caption": "Fig. 3 Schematic of a preloaded three-pad AFB",
        "texts": [
            "asmedigitalcollection.asme.org/ on 10/04/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use where C represents the bearing radial nominal clearance and wj\u00f0h; zj\u00de is the elastic deflection of the underlying support structure, calculated following the independent elastic foundation model presented in Ref. [16]. rp is the hydrodynamic preload of each top foil pad. The hydrodynamic preload is the difference between the nominal clearance C and the minimum clearance Cm, i.e., rp \u00bc C Cm, as illustrated in Fig. 3. The hydrodynamic preload is achieved by making the bearing sleeve with noncircular contour with three lobes (Fig. 3). The hydrodynamic preload results in different clearance along the circumferential direction with larger clearance at the leading and trailing edges. It should be noted that the hydrodynamic preload is different from a mechanical preload which is commonly generated by the manufacturing error of the foils. Typically, overall curvature of the foils (upon following a typical forming/heat treatment process) does not match to the sleeve and foils do not ideally sit on the sleeve. The mechanical preload that comes from the aforementioned manufacturing issue also exists in the single pad foil bearings designed with uniform assembly clearance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000335_j.euromechsol.2012.10.011-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000335_j.euromechsol.2012.10.011-Figure8-1.png",
        "caption": "Fig. 8. Relation between the input compressive force FA and the elastica length variation inside the channel Dl for the case when h \u00bc 0.1, F1 \u00bc 2, mB \u00bc 1.0, and ml \u00bc 0.5. For solid lines, a \u00bc b \u00bc 100. For dashed lines, a \u00bc b \u00bc N.",
        "texts": [
            " At point c, the friction force at clamp B exceeds the maximum static friction, and FA cannot be increased further. For this case Dl\u00f0C\u00demax is the length increment at point d, which is 0.03447. If the pushing force FA decreases gradually, the loadedeflection curve traces points c, d, e, a, and back to the origin. From point c to d, the elastica is in rolling line contact. At point d the contact pattern evolves to rolling point contact. At point e, the rolling point contact evolves to sliding point contact until point a. Apparently, hysteresis occurs within a loadeunload cycle. Fig. 8 shows the effects of spring constants a and b on the loade deflection curves. The solid and dashed lines represent the cases when a \u00bc b \u00bc 100 and a \u00bc b \u00bcN, respectively. Generally speaking, Dl is larger when the spring constants are larger. In the case of line contact, both ends A and B are moment free. Therefore, torsional spring has no effect in line-contact deformation. In the case when the spacing 2h is not large, it is possible to use small-deformation theory to investigate the frictional contact in the constrained elastica"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000513_smc.2014.6974513-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000513_smc.2014.6974513-Figure4-1.png",
        "caption": "Figure 4 The solution of the elbow joint angle",
        "texts": [
            " For the convenience of discussion, we define some vectors as follows: w : A vector from the point S to the point W; e : A vector from the point S to the point E; V : The vector from the point S to the point Q in the reference plane; k p, : The vectors which are perpendicular to w in the B. Resolutions of Joint Angles 1) Determining the Elbow Joint Angle In order to separate the elbow joint angle, we project the arm plane to the plane which is perpendicular to the axis of joint 4. The projected plane is denoted as S E W\u22a5 \u22a5 (see Figure 4). Using the cosine law for the triangle S E W\u22a5 \u22a5 , we can have 2 2 2 4 acos 2 S E E W S W S E E W \u03b8 \u03c0 \u03b2 \u03c0 \u22a5 \u22a5 \u22a5 \u22a5 \u22a5 \u22a5 \u22a5 + \u2212 = \u2212 = \u00b1 (5) where, 22 2 2 2 2 2 4 3 5, ,S W d S E d E W d\u22a5 \u22a5 \u22a5 \u22a5= \u2212 = =w . Given the pose of the end-effector, we can get two resolutions for 4\u03b8 . 2) Determining the Shoulder Joint Angles Given the arm angle \u03c8 , the transformation matrix representing the rotation of angle \u03c8 about the vector w is: ( )20 0 0 3 sin 1 cos\u03c8 \u03c8 \u03c8\u00d7 \u00d7 = + + \u2212 R I w w (6) where, the left superscript \u201c0\u201d indicates the reference frame, 3I is a 3\u00d73 identity matrix; 0 \u00d7w is the skew symmetric matrix of 0 w "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002889_s40436-016-0158-1-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002889_s40436-016-0158-1-Figure5-1.png",
        "caption": "Fig. 5 The thermal deformation vector diagram of the motorized spindle",
        "texts": [
            " It shows that the thermal deformation of the motorized spindle begins from the center of the spindle. The deformation is gradually accumulated, and the deformation of the stator is larger. Front and rear bearings are the parts of high temperature and large deformation in the motorized spindle. The thermal deformation of the ends of the motorized spindle is larger due to the deformation accumulation. The cooling water passage of the motor takes away a portion of the heat, thus the thermal deformation is relatively small. As shown in Fig. 5, the main thermal deformation direction of the stator and the rotor is radial deformation, and the deformation is gradually accumulated from center to edge. The front and rear bearing thermal deformation is shown in Fig. 6. The differences of the front bearings inner ring and the outer ring thermal deformation are around 4 lm and 7 lm, respectively, and the differences of the rear bearings are around 2 lm and 4 lm. The front bearings thermal deformations are larger than the rear bearings both in radial and in axial directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001900_icra.2015.7140018-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001900_icra.2015.7140018-Figure3-1.png",
        "caption": "Fig. 3. (a) Symmetric amplitude changes (\u03b4A) of left and right wing for lift Fz . (b) Asymmetric amplitude changes (\u03b4A) of left and right wing for roll torque Tx. (c) Bias changes (\u03b4\u03c60) of left and right wing for pitch torque Ty . (d) Anti-symmetric split cycle changes (\u03b4\u03c3) of left and right wing for Fx. (e) Symmetric split cycle changes (\u03b4\u03c3) of left and right wing for yaw torque Tz .",
        "texts": [
            ", mg = 1 2\u03c1airCLR 3 w c\u0304r\u0302 2 2(S)\u03c92 wA 2 0, where CL is the mean lift coefficient averaged over one wing stroke [1], and \u03c9w = 2\u03c0f is the wing angular velocity. Sensitivities are defined when kinematic parameters are deviated from their nominal values. Based on the assumption of near-hovering condition and the method in [11] [16], for small deviations from the nominal kinematics parameters in amplitude \u03b4A, bias \u03b4\u03c60, and split cycle \u03b4\u03c3, it can be shown that 1) The lift force Fz is due to symmetric amplitude changes of the left and right wing, i.e., Al = A0 + \u03b4A and Ar = A0 + \u03b4A, as shown in Fig. 3(a), is \u03b4Fz = 1 2\u03c1airCLR 3 w c\u0304r\u0302 2 2(S)\u03c92 wA 2 0 ( 2\u03b4A A0 ) . 2) The roll torque Tx due to asymmetric amplitude changes of the left and right wing, i.e., Al = A0 + \u03b4A and Ar = A0 \u2212 \u03b4A, as shown in Fig. 3(b), is \u03b4Tx = 1 2\u03c1airCLR 3 w c\u0304r\u0302 2 2(S)\u03c92 wA 2 0rcp ( 2\u03b4A A0 ) , where rcp = r\u030233(S) r\u030222(S) Rw is the center of pressure on the wing. 3) The pitch torque Ty due to symmetric bias changes of the left and right wing, i.e., \u03c60l = \u03b4\u03c60 and \u03c60r = \u03b4\u03c60, as shown in Fig. 3(c), is \u03b4Ty = \u2212rcpFzsin(\u03b4\u03c60). 4) The yaw torque Tz cannot be realized by the amplitude and bias change, so the split cycle method introduced in [11] is adopted here to generate yaw torque and longitudinal horizontal force Fx. Specifically, when the left and right wing are anti-symmetric for split cycle, i.e., \u03c3l = \u03c3 and \u03c3r = 1 \u2212 \u03c3, as shown in Fig. 3(d), it can be shown that \u03b4Tz = 1 8\u03c1airCDR 4 w c\u0304r\u0302 3 3(S)\u03c92 wA 2 0 ( 1\u22122\u03c3 \u03c3(1\u2212\u03c3) ) , where \u03c3 = 0.5 \u2212 \u03b4\u03c3, for small \u03b4\u03c3, ( 1\u22122\u03c3 \u03c3(1\u2212\u03c3) ) \u2248 2\u03b4\u03c3 0.25 = 8\u03b4\u03c3. 5) Similarly, longitudinal horizontal force Fx can be generated with symmetric split cycle on the left and right wings, i.e., \u03c3l = \u03c3r = \u03c3 and \u03c3 = 0.5 \u2212 \u03b4\u03c3, as shown in Fig. 3(e): \u03b4Fx = 1 4\u03c1airCDR 3 w c\u0304r\u0302 2 2(S)\u03c92 wA 2 0Cscx ( 1\u22122\u03c3 \u03c3(1\u2212\u03c3) ) , where 4Cscx CD CL \u2248 1.5 and CD CL \u2248 1. From 1)-5) and mg = 1 2\u03c1airCLR 3 w c\u0304r\u0302 2 2(S)\u03c92 wA 2 0, we can derive all the sensitivity functions summarized in the Table I. To illustrate the large value of sensitivities, consider small change of kinematics, for example, \u03b4A A0 = 6deg/60deg = 0.1 and \u03b4\u03c60 = 5.7deg \u2248 0.1. With such small change of kinematics, we have \u03b4Fz = 0.2mg, \u03b4Tx = 0.2mgrcp and \u03b4Tx = \u22120.1mgrcp, i.e. lift has a 20% of variation relative to the body weight, and roll torque and pitch torque all have very large variations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002625_1077546315627242-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002625_1077546315627242-Figure11-1.png",
        "caption": "Figure 11. Details of the faulted gear.",
        "texts": [
            " It weighs about 210 pounds and has a modular design comprising of a two stage oil-lubricated parallel shaft gearbox with rolling bearings, a bearing loader, and a programmable magnetic brake. The gearbox is driven by a variable frequency drive (VFD), controlled by a programmable controller with a display panel showing input speeds and a tachometer showing the r/min of the input shaft. The elements of the DDS are designed such that a variety of configurations can be tested. The manufacturer also supplied chipped (Figure 11) and missing tooth gears to simulate faults in the DDS. The faults were created using a rotary tool. The input shaft has a gear with 32 teeth meshing with an 80-tooth gear on the intermediate shaft. On the output section the 40-tooth gear on the intermediate shaft meshes with the 72-tooth gear on the output shaft. For the present study, a combination of fault and operating conditions (Table 2) were simulated by suitable replacement of a 40-tooth gear with its chipped tooth replica (Figure 11). Tests were run for various steady input speeds starting from 10Hz, 20Hz and 40Hz. The line diagram of the gear meshing is shown in Figure 12. One of the key objectives of this experimental study is to observe the effect of braking torque in the diagnostics of gear faults. The brake used in DDS is of particle magnetic type and has a torque range of 1\u201330 lbf. The braking torque is measured by the voltage reading in voltmeter connected to the brake, which is then converted to the units of torque by suitable calibration factors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002523_j.measurement.2016.02.017-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002523_j.measurement.2016.02.017-Figure13-1.png",
        "caption": "Figure 13 Reaction force vector based on torque analysis on X-axis linear guideways",
        "texts": [
            " Secondly, it should be measured longitudinal and lateral distance of this point until the exact location of the cutting forces at the tip of cutting tool, and distance from the machine bed according to real dimensions of the CNC machine. The values are distances from center of gravity (G). Eq. 4 Eq. 5 Where here; is the total mass of system. The load points of cutting force and the weight force on X-axis are necessary to calculate normal force at contact surfaces. Figure 11 a-c, shows the forces and their distance from the certain points relative to coordinate axes. According to Figure 11 to Figure 13, as a rigid body model, the system of forces can be projected to each point on a body with the corresponding moment. Torque, moment or moment of force, is a force to rotate an object around an axis. Just as a force is a push or a pull, a torque can be thought of as a twist to an object. Mathematically, torque is defined as the cross product of the distance vector and the force vector, which tends to produce rotation. The force vector of cutting force components and weight force creates a torque or couple around the center of gravity. This torque causes reaction forces on the contact surfaces, which is illustrated in Figure 13. The reaction forces increase the friction forces as depicted in Figure 14. The exerted torques and forces can be calculated form the following equations for any component in X, Y and Z directions: From fig. 11 to 13 (a), From fig. 11 to 13 (b), From fig. 11 to 13 (b), From fig. 11 to 13 (c), Because of the forces affecting on the objects in the stationary state. Then: The free body diagram shown in Figure 14. The total friction force in X-axis is as following: Eq. 6 Where; (refer to Figure 8). The static and kinetic coefficient friction ( ) are considered to identify the friction force in X- axis in both of stationary and moving table on the guideways"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003553_s1560354718040081-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003553_s1560354718040081-Figure2-1.png",
        "caption": "Fig. 2. Coordinate systems: a fixed coordinate system Oxy and a moving coordinate system Cx1x2 attached to the body.",
        "texts": [
            " 4 2018 \u2212 from the side of the fluid, the body is acted upon by forces and torques due to added masses, circulation of the velocity of the fluid about the body, and fluid viscosity. We assume that the viscous forces and torques are proportional to the corresponding velocity components of the body (see [10]); \u2212 the system is not acted upon by gravitational forces. Under the assumptions made we consider the problem of investigating the motion regimes of the above system and the possibility of unbounded increase in its kinetic energy. To describe the position and orientation of the body, we define two coordinate systems (see Fig. 2): a fixed coordinate system Oxy and a moving coordinate system Cx1x2 attached to the body. The position of the origin of the moving coordinate system C and the directions of the axes Cx1 and Cx2 are chosen so as to simplify the form of the kinetic energy of the system (see Section 1.2). We specify the position of the body in the coordinate system Oxy by the radius vector r = (x, y) of point C, and the orientation by angle \u03d5 between the positive directions of the axes Ox and Cx1 (see Fig. 2). Thus, the configuration space of the system Q = {q = (x, y, \u03d5)} coincides with the motion group of the plane SE(2). Let v = (v1, v2) denote the velocity vector of point C in the moving coordinate system and let \u03c9 be the angular velocity of the body. Then the following kinematic relations hold: x\u0307 = v1 cos \u03d5 \u2212 v2 sin\u03d5, y\u0307 = v1 sin \u03d5 + v2 cos \u03d5, \u03d5\u0307 = \u03c9. (1.1) 1.2. Equations of Motion The motion of an arbitrary smooth profile in an ideal fluid is described by the Chaplygin equations [16, 21], which generalize the Kirchhoff equations [26]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003644_tr.2018.2870276-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003644_tr.2018.2870276-Figure4-1.png",
        "caption": "Fig. 4. Effects of shock-based failure on the profile of high-speed train wheels.",
        "texts": [
            " Therefore, both wear and external shock should be considered during the optimization of the reprofiling policy for high-speed train wheels. For the purpose of describing the influence of the shock process on the wheel, we define W (t) as the magnitude of an external shock at time t. In Fig. 3, the wheel experiences shock-based failure when the magnitude of the shock W5 at time t5 is greater than the critical level \u03c9 of an extreme shock (W5 > \u03c9). The shock-based failure can be removed by cutting material off the wheel. The result, as shown in Fig. 4, is that the diameter suddenly decreases. In the reprofiling problem, we consider shock-based failures by linking the impact of the random shocks to the degradation level. The level of wheel flange thickness recovery is approximately proportional to the loss in the diameter due to reprofiling, as shown in (1). The lifespan of a wheel is controlled by its minimum diameter. On one hand, the more the flange thickness is recovered by reprofiling, the smaller the diameter becomes, and the earlier the wheel should be replaced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000221_acc.2012.6315545-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000221_acc.2012.6315545-Figure1-1.png",
        "caption": "Fig. 1. The camera\u2019s view: projection of a road with longitudinal axis xr .",
        "texts": [
            " Finally, the quad-rotor is required to land autonomously in a position near the end of the road. The main goal is to design a road following strategy based on computer vision and switching control, with the purpose of stabilizing the vehicle\u2019s altitude, heading angle and lateral position (z, \u03c8 and y states respectively) w.r.t. the road, while traveling at constant forward speed (x\u0307 = c) in presence of external disturbances. Consider that the vehicle is flying over a road having smooth curves, which is represented in the camera\u2019s image as a group of straight and parallel lines, see Fig. 1. A straight line in the image can be seen as a segment of infinite length and whose center of gravity belongs to the straight line [9]. Based on the Hough transform method for line detection, a straight line is represented as \u03c1 = x cos \u03b8 + y sin \u03b8. The center of gravity (xg, yg) of each straight line detected can be computed as xig = cos(\u03b8)\u03c1 and yig = sin(\u03b8)\u03c1, where the super-index i stands for the line i. It is possible to assign initial and final bounds to the line. Let\u2019s define (xiI , y i I) as the initial point of the line, located in a place below the image\u2019s lower margin, and let (xiF , y i F ) be the final point of the line, located in a place above the image\u2019s upper margin",
            " This average line will uniquely represent the road in the image with a simple pair of initial and final coordinates ( xI = xiI i , yI = yiI i ) ; ( xF = xiF i , yF = yiF i ) (3) where i is the number of lines grouped together, (xI , yI) represents the initial (lowermost) road coordinate and (xF , yF ) represents the final (uppermost) road coordinate. The angle \u03c8r between the camera\u2019s xc axis and the point given by (xF , yF ) can be computed using (xI , yI) and (xF , yF ) in the two argument function atan2 as \u03c8r = atan2(yF \u2212 yI , xF \u2212 xI). The angle \u03c8r is used for obtaining the desired heading angle \u03c8d that will align the vehicle\u2019s xaxis (xh) with the road\u2019s longitudinal axis (xr), see Fig. 1. \u03c8d will finally be expressed as \u03c8d = \u03c8r+ \u03c0 2 , where the therm \u03c0 2 is added to adjust \u03c8d to a value of zero when \u03c8r is aligned vertically with xh. Consider an image-based distance ecx along the camera\u2019s xc axis. ecx is defined between the road\u2019s center of gravity projection (xg, yg) and the vehicle\u2019s center of gravity projection (x0, y0), see Fig. 1. In the case when xI > xF , one has ecx = ( xI \u2212 xF 2 + xF ) \u2212 cw 2 (4) where cw represents the width of the image in pixels. In the case when xI < xF , xI must be replaced by xF and viceversa. The lateral position of the aerial vehicle w.r.t. the road can be estimated from ecx as ey = z ecx \u03b1x (5) where z represents the distance existing between the helicopter and the road (altitude), and \u03b1x represents the camera\u2019s focal length in terms of pixel dimensions in the xc direction. Consider a camera-vehicle arrangement moving in a 3- dimensional space with respect to a rigid scene",
            " Following a similar approach than the one presented in [14], it is possible to find a common Lyapunov function for the closed-loop system formed from applying the two controllers of the lateral position dynamics. However, working in this way, the same pole locations have to be chosen for both cases, which in fact is not the case (different gain values are being applied). Let\u2019s define dc as the distance measured from the vehicle\u2019s center of gravity projection to the point where the camera loses the image of the road, see Fig. 1. Thus, a change of coordinates can be made such that yd1 = x3 + dc and y\u0307d1 = yd2 = x4 is its derivative. From equation (17) one has y\u0307d1 = yd2 , y\u0307d2 = \u2212kL3yd1 \u2212 kL4yd2 \u2212 kLI\u03be, \u03be\u0307 = \u2212yd1 , with eyd = (yd1 , yd2 , \u03be)T . It can be defined a state-dependent switched linear system, given by the closed-loop system together with the switching conditions e\u0307yd = { ALeyd if |yd1 | < 0 ANLeyd if |yd1 | \u2265 0 (20) It is clear that each individual system in equation (20) is stable, since the matrices AL and ANL are Hurwitz"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002018_0305215x.2014.905551-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002018_0305215x.2014.905551-Figure1-1.png",
        "caption": "Figure 1. Circumferential stress distributions for the cases of (a) wavy edge defect and (b) centre buckle defect.",
        "texts": [
            "kr \u00a9 2014 Taylor & Francis D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 8: 54 3 0 N ov em be r 20 14 2 N. Pholdee et al. the coiling process was revisited by Park et al. (2014) to include the effect of the crown shape of the spool in the analysis. In summary, manufacturing a flat strip without the crown shape is the best solution for preventing flatness defects on the metal strip during the strip coiling process. However, it is almost impossible to make flat strips in industry. Decreasing the strip crown or increasing its exponent, as shown in Figure 1, is difficult to achieve because of production constraints.Therefore, processing parameters such as spool geometry and coiling tension were selected as controlling parameters to improve the strip flatness during the strip coiling process in the present work. To determine the optimal values of these processing parameters, various optimization studies are necessary. In general, optimization algorithms can be classified into two categories: those with a function derivative, also known as gradient-based methods; and those that do not apply a function derivative, for example evolutionary algorithms (EAs)",
            " For this purpose, the optimal processing parameters of spool geometry and coiling tension were determined by introducing the objective function to minimize axial inhomogeneity of the stress distribution and the maximum stress, which may cause the irregular surface profile of the strip during the coiling process. During the strip coiling process, a wavy edge will form when the circumferential stress at the middle zone of the strip is highly compressed while the two edges are tensile or lightly compressed. On other hand, if the middle zone of the strip is highly tensile while the two edges are compressed or lightly tensile, a centre buckle will occur, according to previous studies (Yanagi et al. 1998; Park et al. 2014). Figure 1(a) and (b) shows the circumferential stress \u03c3\u03b8 distributions along the z direction within a thin strip that incurred a wavy edge and a centre buckle, respectively. In the strip coiling process, when the strips are coiled, the middle zone (z = 0) of the strip at the inner coil will be highly compressed compared with the two edges owing to the coiling tension and strip crown. In this case, the centre buckle at the inner coil is unlikely and the wavy edge is likely to form. Therefore, the wavy edge defect at the inner coil becomes a major issue during the strip coiling process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000945_iceecs.2014.7045248-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000945_iceecs.2014.7045248-Figure1-1.png",
        "caption": "Fig. 1. Cross section of a BLDC motor",
        "texts": [
            " Cogging torque is an unwanted phenomenon in electrical machines, as a product of reluctance variations between stator tooth and rotor magnetic poles [1]. As slotless BLDC motors also have many disadvantages, thus the characteristic of slotted and slotless BLDC motors must be described to choose the best type of BLDC motors for certain applications. This research was done to study the characteristics of slotted and slotless BLDC motor by using 2D finite element simulation. II. BRUSHLESS DC MOTOR Brushless DC motor is a DC motor without brush component. The switching of polarization is done by using electrical switching. Fig 1 shows the cross section of a BLDC motor [2]. Permanent magnets are used as moving parts and stator is the static parts. The nonlinear magneto-static field in an electric machine can be defined by the following Maxwell\u2019s equations [3] : \u00d7 H = J (1) B = \u00b5H (2) The magnetic vector potential A is defined in equation 3. B = \u00d7 A (3) where A is the axial component of the magnetic vector potential, \u03bd is the reluctivity of the material and J is the current density. The mathematical model of a BLDC motor prototype can be represented as:[4] (4) where \u03bd is the reluctivity of the material, J is the applied source current density, Jm is the equivalent current density of the permanent magnet, \u03c9 is the angular frequency, and \u03c3 is the conductivity of the material"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure2.3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure2.3-1.png",
        "caption": "Fig. 2.3 Conditional lines of the leakage fluxes and fluxes of self- and mutual induction of the induction machine windings (a) lines of the leakage fluxes and fluxes of self- and mutual induction of the stator winding; (b) lines of the leakage fluxes and fluxes of self- and mutual induction of the rotor winding; (c) lines of the leakage fluxes and fluxes of self- and mutual induction of the stator and rotor windings",
        "texts": [
            "7) for the stator and rotor windings cannot be used in practice. For the total fluxes of self-induction of the stator winding \u03a61 and of the rotor winding \u03a62, we can use the following statements arising from obvious physical considerations. The flux \u03a61 does not fully cover the rotor winding, and the flux \u03a62 does not fully engage with the stator winding, which is associated with the leakage phenomenon that occurs in an electric machine. To take this into account, we distinguish the leakage flux of the stator winding \u03a61\u03c3 covering only this winding (Fig. 2.3a) and the leakage flux of the rotor winding \u03a62\u03c3 associated only with this 2.1 An Induction Machine with a Single-Winding Rotor: Voltage Equations. . . 17 winding (Fig.2.3b). For these fluxes, on the basis of (2.3), we can obtain the obvious expressions \u03a61\u03c3 \u00bc \u03a61 \u03a612 \u03a62\u03c3 \u00bc \u03a62 \u03a621 \u00f02:9a\u00de Here, the flux\u03a612 is created by the current in the stator winding. It represents the flux of mutual induction, and therefore it crosses both the stator and rotor windings. The designation of flux \u03a612 indicates that it crosses, in this case, the rotor winding. In order to emphasize that flux \u03a612 is also engaged with the stator winding, we introduce a new designation in the form \u03a611\u00bc\u03a612",
            " Taking into account the conditions \u03a611\u00bc\u03a612 and \u03a622\u00bc\u03a621, the equations (2.9\u0430) obtain \u03a61\u03c3 \u00bc \u03a61 \u03a611 \u03a62\u03c3 \u00bc \u03a62 \u03a622 \u00f02:9b\u00de From here, for the total flux of self-induction of the stator winding \u03a61, it follows that \u03a61\u00bc\u03a611 +\u03a61\u03c3, and for the total flux of self-induction of the rotor winding \u03a62, we have \u03a62\u00bc\u03a622 +\u03a62\u03c3. Using these conditions, equations (2.3) for the resulting fluxes \u03a61\u0440 and \u03a62\u0440 receive the form \u03a61\u0440 \u00bc \u03a61 \u00fe \u03a621 \u00bc \u03a61\u03c3 \u00fe \u03a611 \u00fe \u03a621 \u03a62\u0440 \u00bc \u03a62 \u00fe \u03a612 \u00bc \u03a62\u03c3 \u00fe \u03a622 \u00fe \u03a612 \u00f02:10\u00de In (2.10), the fluxes \u03a611 and \u03a612 have the same values by definition, i.e., \u03a611\u00bc\u03a612 (Fig. 2.3a). Analogously, for the fluxes \u03a622 and \u03a621, we obtain \u03a622\u00bc\u03a621 (Fig.2.3b). In this regard, equations (2.10) can be represented in the following form: \u03a61\u0440 \u00bc \u03a61\u03c3 \u00fe \u03a611 \u00fe \u03a621 \u00bc \u03a61\u03c3 \u00fe \u03a6m \u03a62\u0440 \u00bc \u03a62\u03c3 \u00fe \u03a622 \u00fe \u03a612 \u00bc \u03a62\u03c3 \u00fe \u03a6m \u00f02:11\u00de where \u03a6m\u00bc\u03a611 +\u03a621\u00bc\u03a622 +\u03a612. By (2.11), the resulting fluxes \u03a61\u0440 and \u03a62\u0440 are represented as the sum of the leakage fluxes \u03a61\u03c3 and \u03a62\u03c3 and the resulting flux of mutual induction \u03a6m. The conventional picture of the magnetic fluxes corresponding to equations (2.11) is shown in Fig. 2.3c. 18 2 Fundamentals of the Field Decomposition Principle The leakage fluxes\u03a61\u03c3 and\u03a62\u03c3 induce the leakage emfs E1\u03c3 and E2\u03c3 in the stator and rotor windings. In order to determine emf E1\u03c3 and E2\u03c3, it is necessary to calculate the magnetic flux linkages \u03c81\u03c3 and \u03c82\u03c3 produced by the leakage fluxes \u03a61\u03c3 and \u03a62\u03c3. The flux linkages \u03c81\u03c3 and \u03c82\u03c3 can be calculated by expressions in the form [9] \u03c81\u03c3 \u00bc 2\u03bc0 w2 1 pq1 lz1\u03bb1 ffiffiffi 2 p I1 and \u03c82\u03c3 \u00bc 2\u03bc0 w2 2 pq2 lz2\u03bb2 ffiffiffi 2 p I2 \u00f02:12a\u00de In the case of the squirrel-cage rotor, the leakage flux linkage \u03c82\u03c3 is determined as \u03c82\u03c3 \u00bc \u03bc0lz2\u03bb2 ffiffiffi 2 p I2 \u00f02:12b\u00de In (2",
            "19), the emfs E1m and E2m are determined as a result of the sum of the emfs E11,E21 and E22,E12, respectively. The flux linkages \u03c811,\u03c821,\u03c822 and \u03c812 used in (2.19) arise from the expressions \u03c811 \u00bc w1kw1\u03a611;\u03c821 \u00bc w1kw1\u03a621 \u03c822 \u00bc w2kw2\u03a622;\u03c812 \u00bc w2kw2\u03a612 \u00f02:20\u00de As follows from (2.20), the flux linkages \u03c811, \u03c812, \u03c822 and \u03c821 are caused by fluxes \u03a611\u00bc\u03a612 and \u03a622\u00bc\u03a621. To determine these fluxes, we use the provision that the fluxes of self- and mutual induction \u03a611\u00bc\u03a612 and \u03a622\u00bc\u03a621 have the same picture of distribution in consideration of a symmetrical electric machine (in Fig. 2.3a, b) the configurations of the magnetic circuits are the same for 20 2 Fundamentals of the Field Decomposition Principle the fluxes of self- and mutual induction \u03a611\u00bc\u03a612 and \u03a622\u00bc\u03a621. The permeance factor for these fluxes is determined as \u03bbm\u00bc \u03bc0/\u03b4 kHk\u03b4,, where kH is the saturation factor, k\u03b4 is the air gap factor, and \u03b4 is the length of the air gap. The magnetic flux density in the air gap stimulated by the currents in the stator and rotor windings is calculated by the expression Bk \u00bc \u03bbmFk \u00bc \u03bc0 kH\u03b4 0 mkwkkwk \u03c0p ffiffiffi 2 p Ik, where k\u00bc 1, 2",
            " According to equations (2.24), an electric machine operates like a transformer. Therefore, no-load conditions of an electric machine can be implemented when the value of the rotor winding current is equal to zero, i.e., I2\u00bc 0. Under these conditions, the no-load current Im is flowing in the stator winding. The current Im creates the mmf of the stator winding Fm equal to Fm \u00bc m1w1kw1 \u03c0p ffiffiffi 2 p Im. The mmf Fm produces the magnetic field, creating in an electric machine the flux of mutual induction\u03a611\u00bc\u03a612 (Fig. 2.3). The condition\u03a611\u00bc\u03a612 means that the same flux of mutual induction created by the stator winding current, in the first case, crosses the stator winding, and in the second case, couples with the rotor winding. This flux magnetizes the main magnetic circuit of an electric machine. Therefore, it reflects the magnetizing flux of an electric machine \u03a6m determined as \u03a6m\u00bc\u03a611\u00bc\u03a612. In this connection, the no-load current Im producing the magnetizing flux \u03a6m represents the magnetizing current of an electric machine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000536_worv.2013.6521918-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000536_worv.2013.6521918-Figure7-1.png",
        "caption": "Figure 7. Scooping along osculating circle.",
        "texts": [
            " In this section, the scooping procedure for the non-rigid and loaf shape food such as tofu (soybean curd) is described. In this paper the proposed method described in section 3.1\u20133.3 is called ordinary scooping. Also the proposed method described in section 3.4 is called remnant scooping. Before executing food scooping, setting up of three parameters is required. Two parameters are related to the size of scooping food, the width and the height. The depth of scooping is calculated according to the height of food on plate. The last parameter is the diameter of a circle which defines trajectory of the spoon (see Fig. 7). Food can be scooped up by moving the spoon along that circle osculating the spoon. In order to realize successful scooping, determination of a scooping position of the food is very important. The fol- lowing is the proposed algorithm to determine a scooping position when the food is put on a shallow plate. 1. The measured 3D points of the food are divided into three groups depending on their heights (z coordinates), where the xyz coordinate system is set so that the x-y plane is parallel to the horizontal surface of the table",
            " By back projection of point E to the top surface point in 3D, this 3D point is determined as the scooping position of the food. In the case of a second or subsequent scooping, determination of a scooping position is realized by the same algorithm as the first scooping (Fig. 6). Scooping of the food with the spoon is executed along the osculating circle of spoon. Spoon direction on the x-y plane of manipulator coordinate system is parallel to vector M given in section 3.2. The spoon edge is inserted from point E and rotate the spoon around the center of the osculating circle as shown in Fig. 7. The velocity of scooping motion is kept constant. The scooping method mentioned above cannot scoop up a food completely. In order to scoop the remnant food, an additional procedure with a longer stroke along the plate is introduced. Note here that it is more suitable to use a plate whose slope of edge is larger for remnant scooping. In other words, flat plates are not preferred for the scooping. The following is an additional procedure to solve this problem. First, a curved surface equation is fit to the plate model already known so that a normal vector of arbitrary position of the plate can be calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002067_s11249-013-0291-y-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002067_s11249-013-0291-y-Figure4-1.png",
        "caption": "Fig. 4 Electro-Magnetic Loading Device",
        "texts": [
            " Therefore, a controllable seal system can be established through adjusting the closing force in accordance with the feedback of the measured friction torque and temperature. The electromagnetic loading device of the seal closing force and the testing system of the friction torque, the key part of the entire controllable seal system, are critical to whether the seal controllability can be achieved. Therefore, the design, calibration and installation of the two will be elaborated in the following. The EMLD (Fig. 4) is designed to generate an adjustable axial electromagnetic force, which is an important part of the seal closing force. The other parts of the seal closing force include the hydrostatic force due to the pressurized fluid and the spring force. It should be noted that the secondary O-ring seal between the stator and the seal chamber cannot affect the axial closing force which is orientated toward the axial direction. Therefore, the O-ring force can be neglected. And it also has a negligible effect on the face friction torque transferring from the mating faces to the EMLD, as explained in Sect"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003697_j.ijthermalsci.2018.11.027-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003697_j.ijthermalsci.2018.11.027-Figure11-1.png",
        "caption": "Fig. 11. Thermo-deformation finite element simulation results of thermal balance of spindle.",
        "texts": [
            " From the simulation results, the maximum temperature of the bearing-spindle system is 43.05 \u00b0C when the system reaches a balance, and the maximum temperature occurs at the contact points between the inner raceway and balls. Based on the simulation model of bearing-spindle temperature field, the thermo-mechanical coupling analysis was carried out to calculate the system thermal elongation. The simulation results of the bearingspindle system were post-processed by importing the temperature data into the spindle thermal elongation model. Fig. 11(a) shows thermal distortion of the spindle system after it reaches thermal balance at 10000 r/min. By using the traditional empirical thermal convection parameters, when the steady state is reached, the maximum thermal elongation of the spindle right and left end face (X direction) are 15.34 \u03bcm and 12.34 \u03bcm respectively. When establishing the simulation model of the spindle thermal elongation, the left and right end faces of the spindle were not fixed and were in a free state. Therefore, the actual thermal elongation of the spindle was the sum of two end faces. At the speed of 10000 r/min, the axial thermal elongation can be obtained by Eq. (10): = + = + =+ \u2212L L L \u03bcm\u0394 15.34 12.34 27.68 (10) Based on the optimized thermal coefficients by multi-objective optimization, the corresponding thermal convection parameters are substituted into the bearing-spindle thermal elongation model to obtain the modified thermal elongation in the axial directions. As shown in Fig. 11(b), the maximum thermal elongation of the spindle's right and left end face (X direction) are 12.59 \u03bcm and 9.99 \u03bcm respectively, and the total thermal elongation in the axial direction is: \u2032 = + = + =+ \u2212L L L \u03bcm\u0394 12.59 9.99 22.58 (11) 5.2. Experimental verification of bearing-spindle thermal elongation The test bench of high-precision bearing-spindle system was set up, whose structure is the same as shown in Fig. 10. The high precision laser displacement sensor was fixed on the surface of the test bed with the magnetic table"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001846_978-90-481-9707-1_116-Figure11.4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001846_978-90-481-9707-1_116-Figure11.4-1.png",
        "caption": "Fig. 11.4 Mechanical structure layout and size of the quadrotor",
        "texts": [
            " Quadrotor arms: A protection scheme is developed as a shell to encapsulate the motor into a tight chamber to fix its position. Carbon fiber beam can be mounted and screwed to the side, as shown in Fig. 11.3a. 2. Main body: The main frame structure has four slots for the quadrotor arms and contains two layers, where the avionic system is placed on the top and the battery is located in the lower level, as shown in Fig. 11.3b. It is fabricated with acrylonitrile butadiene styrene (ABS). The mechanical structure layout of the proposed quadrotor MAV is shown in Fig. 11.4a with all the parts assembled together as shown in Fig. 11.4b. Distance of two diagonal rotors is 142.54 mm, with a total height of 22.50 mm. Details of weight breakdown are reflected in Table 11.5. In this table, the estimated weight is approximated based on the design guideline shown in this chapter, while the current weight on the right of the table is the measured weight of the quadrotor MAV prototype, code name KayLion, made by National University of Singapore, following the guideline. Hardware system and the avionics are the core of any unmanned aircraft design"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000471_iros.2012.6385571-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000471_iros.2012.6385571-Figure2-1.png",
        "caption": "Fig. 2: Overview of pneumatic soft actuator",
        "texts": [
            " This length direction force, hereinafter an axial force, can reduce a human load by being converted to an assist torque according to a moment arm, which is a length between a fixed point of actuator and human join. On the other hand, a human joint is also compressed by this axial force. It is desirable that an actuator applies a vertical assist force to a human body in order to protect a knee 978-1-4673-1736-8/12/S31.00 \u00a92012 IEEE 1239 joint from an excessive axial force. The pneumatic soft actuator shown in Fig.2 (a)(b) consists of the expansion unit, nylon bands. The expansion unit has four aluminum film balloons covered with a nylon band, and has 85mm in width, 80mm in depth and 0.6mm in thickness, respectively. Aluminum film balloon shown in Fig.2 (c) is made by gluing outsides of two aluminum films. Therefore, this balloon can have 70 mm in inner width, 65 mm in inner depth. The lengths of nylon band at thigh and calf are 380mm and 280mm, respectively. When a compressed air is supplied into the balloons, the balloon as shown in Fig.2(c) expands to a height direction. Therefore, the expansion unit also expands to height direction. The expansion force is applied along the length direction of nylon band. The vertical assist force can be increased as shown in Fig.3 since the angle \u03b8 between the nylon band and the human body is increased by fixing the nylon band on back sides of a thigh and calf. The bottom of trousers is fixed to foot by the nylon band to prevent moving the bottom upward according to the axial force. The bottom of trousers is fixed at the heel for converting this axial force to the assist torque for plantar flexion at ankle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure7-1.png",
        "caption": "Fig. 7. Geometric relation for calculation of the separation distance of a tooth.",
        "texts": [
            " However, only exact involute tooth profile, the fixed supported sun gear and error-free conditions are considered in the paper. Some important equations on planetary gear meshing will be summarized in the section for better understanding, more details can be found in the reference [29] . Another essential component in the proposed LTCA model for the three-dimensional simulation of distributed contact stresses is the two-dimensional separation distances between each two engaged tooth flanks. The calculation can be derived from the geometrical relation illustrated in Fig. 7 . If a flanks engages another on point M, the distances h of any point Y on the flank to the tangential plane is to be determined. The location of point Y is defined with a specific length l to point M along the tangent t in the transverse plane. The given length l is equal to the inner product of the vectors r YM ( = r Y \u2212 r M ) and t , i.e., [ r Y ( \u03beY ) \u2212 r M ( \u03beM )] \u00b7 t ( \u03beM ) = l. (33) The variable \u03beY for the point Y can be thus solved. The corresponding separation h can be calculated with substituting the variable \u03beY into the equation: h = [ r Y ( \u03beY ) \u2212 r M ( \u03beM )] \u00b7 n ( \u03beM ) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002948_icems.2019.8922014-Figure15-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002948_icems.2019.8922014-Figure15-1.png",
        "caption": "Fig. 15. Air flow comparison",
        "texts": [
            " In this paper, the losses at low to middle speed area and low to middle torque area are investigated because these driving points are frequently used in the city driving. Fig. 14 shows the analysis results of eddy current loss, windage loss, torque, output power, and efficiency. Calculation method for efficiency is expressed as Pout\u03b7 W W WPout c i w     (3) where Ww is windage loss. As shown in Fig. 14, it is revealed that eddy current loss and windage loss is reduced by dividing the rotor of the proposed SRM. Also, torque and output power are little decreased result from deterioration of stacking factor of the rotor core by the inserted shrouds. Fig. 15 shows air flow comparison of the conventional rotor and the rotor divided by 20. As shown in Fig. 14 and Fig. 15, windage loss of the conventional rotor structure is increased because air flows between the salient poles and the rotor sweep out air. Unlike conventional result, windage loss of the proposed rotor structure is reduced because air flows in each divided core and air rotates with the rotor. Finally, the number of division of the rotor core is optimized as shown in Fig. 16. As shown in Fig. 16, 10 division of the rotor core is the best at low speed and low to middle torque area because the effect of eddy current loss is higher than windage loss"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003603_1350650118800581-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003603_1350650118800581-Figure6-1.png",
        "caption": "Figure 6. The TC rotor-bearing model: (a) CAD model; and (b) FEM model.",
        "texts": [
            " As can be seen from the following, this work systematically reveals the influence of the reduced load capacities and bearing torques caused by the circumferential or/and axial grooves on the investigated TC rotor\u2013FRB system subsynchronous oscillations. In the performed simulations, some intriguing and distinct phenomena, which have not been reported previously, will occur. Rotor-bearing model descriptions and pertinent parameters The TC rotor consists of a shaft, a compressor disk, and a turbine disk. The shaft is supported by two fullfloating ring bearings (see Figure 6(a)). In order to carry out numerical simulations, the continuous TC rotor is modeled by three Timoshenko beam elements of constant cross-section and discretized to four nodes (see Figure 6(b)). The compressor and turbine disk are considered as rigid bodies, of which the properties are concentrated to the corresponding nodes of the discretized rotor model. Based upon the finite element theory and the MATLAB toolbox Rotor Software V1 described in Friswell,25 the governing equations of motion for the TC rotor are formulated as M\u00bd \u20acq \u00fe C\u00fe jG _q \u00fe K\u00bd q \u00bc Fif g \u00fe Fubf g Wf g \u00f014\u00de The differential equation of motion given by equation (14) takes into account the inner fluid film force vector {Fi}, the unbalanced force vector {Fub}, the static gravitational force vector {W} in Y direction only, and the generalized coordinates of the flexible rotor {q}\u00bc {x1, y1, 1, 1, x2, y2, 2, 2, x3, y3, 3, 3, x4, y4, 4, 4} T, which consider all the degrees of freedom of the rotating system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001461_j.mechmachtheory.2018.05.013-Figure7-1.png",
        "caption": "Fig. 7. Maneuvering for traversing a point elusive barrier.",
        "texts": [
            " For a given task point Q = (t 1 , t 2 ) = (0 , 0) = O, the self-motion manifolds in the 3D joint space ( \u03b81 , \u03b82 , \u03b83 ) are two circles obtained as the intersection between cylinder { (\u03b81 \u2212 t 1 ) 2 + (\u03b82 \u2212 t 2 ) 2 = d 2 } and two planes { \u03b83 = K 1 } and { \u03b83 = K 2 } , where K 1 and K 2 are the two solutions of \u03b83 in [ \u2212\u03c0, \u03c0 ] that satisfy the following equation: (t 1 \u2212 cos \u03b83 ) 2 + (t 2 \u2212 sin \u03b83 ) 2 = 1 (8) Fixed joint O = (0 , 0) in this example is an interior barrier of the workspace, as justified next. Assume that joint Q coincides with O , as illustrated in Fig. 7 (a). In that case, the robot can only generate task velocities in the direction perpendicular to link OP . If one wishes to move joint Q along the direction L of link OP (for describing a linear trajectory across O , for example), links OP and PQ must be rotated first an angle of 90 \u00b0, so that they remain perpendicular to L ( Fig. 7 (b)). After that, velocities or displacements along L can be generated ( Fig. 7 (c)). This occurs for any orientation \u03b83 of link OP , since when Q coincides with O , link OP can freely rotate without modifying the position of Q (this is because links OP and PQ have the same length). This can also be observed in Eq. (8) , which is satisfied \u2200 \u03b83 when Q = (t 1 , t 2 ) = (0 , 0) = O . Since \u03b83 is not constrained, the self-motion manifolds at O are not two circles but a single cylinder defined by { \u03b82 1 + \u03b82 2 = d 2 } . Note that self-motion manifolds do not vanish at this special barrier, but their dimension instantaneously changes: two circles, which are one-dimensional manifolds obtained by intersecting a cylinder with two planes, transform into the complete cylinder, which is a two-dimensional manifold"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001568_j.mechmachtheory.2018.07.009-FigureC.14-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001568_j.mechmachtheory.2018.07.009-FigureC.14-1.png",
        "caption": "Fig. C.14. A path passing through the acnode of the singularity-manifold in the \u03b81 \u2212\u03b82 space. (a) Path intersecting the singularity manifold which has reduced to an acnode; (b) The motion of the manipulator at the corresponding configuration-space singularity.",
        "texts": [
            ", the angular motion of link l 1 changes from CW to CCW direction and the angular motion of the link l 2 changes from CCW to CW). This motion is responsible for the occurrence of a cusp on the path (see Fig. C.13 a), at the point at which it intersects the singularity manifold. C.3. Double point on the gain-type singularity manifold (acnode) The link dimensions as well as the trajectory tracked by the manipulator are obtained from the singular case in the configuration-space forward dynamics example (refer to Section 7.1 ). The corresponding path in the \u03b81 \u2212\u03b82 space is plotted in Fig. C.14 a. The pose of the manipulator at the configuration-space singularity, where all the joints are collinear, is shown in Fig. C.14 b. It is interesting to note that the singularity manifold for the specified link dimensions degenerates to an isolated point, namely, an acnode. At that point the kinematic consistency equation is identically satisfied for all values of \u02d9 \u03b8, since the coefficients \u03b11 and \u03b12 in Eq. (45) turn out to be zeros. Therefore, the actuators instantaneously regain their independence in the velocity-space at this configuration. This section illustrates the three methods described in Section 2.2.1 for the computation of the actuator forces in the inverse dynamics problem"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002887_fie.2016.7757670-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002887_fie.2016.7757670-Figure1-1.png",
        "caption": "Figure 1: Tasks in simple gear train laboratory exercise",
        "texts": [
            " Tasks for simple gear train laboratory and planetary gear train laboratory In the simple gear train laboratory exercise, the students are required to design a simple gear train with specified speed ratio and output direction. The virtual laboratory provides mechanical parts, such as shaft, shaft holder, gears with different number of teeth, etc., for them to use. Since this is their first virtual laboratory exercise, the students must run through a short tutorial that teaches them how to use the virtual laboratory. Figure 1 shows the tasks for the students in the simple gear train laboratory exercise. The students must first select and mount a gear to a shaft, then mount the shaft on the shaft holders, and adjust the center distance of different gears to form a simple gear train. The students are expected to enhance their understanding of the key concepts of gears and gear trains, such as module, pitch diameter, and idler gears, etc., during this laboratory exercise. After the students have completed the simple gear train laboratory exercise, a planetary gear train laboratory exercise is administered which involves fewer operations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000048_j.jelechem.2013.08.030-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000048_j.jelechem.2013.08.030-Figure1-1.png",
        "caption": "Fig. 1. Construction of thin-layer flow cell: (1) polyacrylic block, (2) Ag/AgCl film, (3) internal flow path (2 mm 22 mm), (4) PTFE spacer (50 lm thickness), (5) PTFE porous membrane containing NPOE (30 lm thickness), (6) conducting polymercoated Pt plate, and (7) PDMS.",
        "texts": [
            "1 M BTPPATFPB, by oxidizing whole of the PEDOT at 0.6 V, and then by reducing it partially at 0.4 V. Before the measurements, the obtained PEDOT-Pt was dipped in NPOE containing 10 3 M of BTPPATFPB for at least 1 h. In the present thin-layer flow cell, NPOE was employed as the Org phase. NPOE has low vapor pressure, high viscosity (13.8 mPa s) [21], and low solubility in water (2.01 lmol dm 3) [22]. The thin layer of NPOE was stable and durable in the flow system. The thin-layer flow cell is shown in Fig. 1. An Ag film (20 mm 10 mm, thickness of 20 lm) coated with AgCl and the 50% oxidized PEDOT-Pt were employed as electrodes for the W and NPOE, respectively, which served as both reference and counter electrodes. The thin-layer flow cell was composed of an upper Ag/AgCl electrode part, a middle part of polytetrafluoroethylene, PTFE, spacer, and a lower PEDOT-Pt part. The Ag/AgCl electrode and the PEDOT-Pt were supported by polyacrylic blocks (20 mm 10 mm, 5 mm height) in order to maintain the flat plane of the electrode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002320_j.proeng.2014.11.647-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002320_j.proeng.2014.11.647-Figure2-1.png",
        "caption": "Fig. 2. Schematic of the DNA-LAPS structure as well as multi-spot and scanning-beam setup.",
        "texts": [
            " A schematic diagram of different strategies for ssDNA immobilization and hybridization on the LAPS surface is shown in Fig. 1. For DNA hybridization, two protocols were used: For immobilization strategy (a) and (b), 150 \u00b5L solution of complementary target ssDNA (1 \u00b5M DNA diluted in 0.1 M PBS, 0.9 M NaCl, pH 7.0) was applied to the chip surface for 15 min. For strategy (c), the same concentration of target DNA was diluted in 1x Tris-EDTA buffer, pH 8.0 and the incubation time was increased up to 45 min. Fig. 2 shows a schematic structure and measurement setup of the DNA-LAPS. The LAPS signal was recorded after each surface functionalization step by means of photocurrent-voltage measurements. The measurements were carried out in the solution of 0.2 mM PBS, pH 7.0 (for immobilization technique (a) and (b)) or 10 mM NaCl, pH 5.4 (for immobilization technique (c)). (c) layer-by-layer adsorption of negatively charged ssDNA on a positively charged PAH. All ssDNA sequences used in this study contain 20 bases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003422_s12206-018-0307-5-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003422_s12206-018-0307-5-Figure3-1.png",
        "caption": "Fig. 3. Testing pneumatic tire.",
        "texts": [
            ", Ltd in Thailand which made of the natural rubber. It had been constructed by enveloping with three layers of the natural rubber components. The solid tire brand KOMACHI which is selected to use with the baggage towing tractors is shown in Fig. 2. The pneumatic tire was regular usage with the baggage towing tractors of the Thai Airways International Public Company Limited. It had the outside diameter and width which were approximate as the dimension of KOMACHI solid tire. The pneumatic tire weight was 9.0 kg. Fig. 3 shows the pneumatic tire which was carried out to determine the vibration effects using the drum testing. The specification of KOMACHI solid tires and pneumatic tires are described in Table 1. The width and rim size of the solid and pneumatic tire were 6.00 and 9 inch (6.00-9), respectively. The steel wheel with a diameter and the E collar profile width of 9 and 4.00 inch (4.00E-9) was proper to fit with the both tires. The weight of both tires was different. The pneumatic tire weight was less than the solid tire 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000842_tmag.2015.2480546-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000842_tmag.2015.2480546-Figure2-1.png",
        "caption": "Fig. 2. Two-pole generator no-load field distribution (healthy operation).",
        "texts": [
            ", EFCAD, was developed by us, and the techniques to consider movement were implemented in this system [10]. The numerical techniques necessary to obtain consistent results are explained in section III. The 10 kVA two-pole generator has a 380 V three-phase wye connected stator and is able to operate at 50 and 60 Hz frequencies. It is moved by means of a dc motor. Fig. 1 shows the bench. As the commercial frequency in Brazil is 60 Hz, the 50 Hz operation was chosen to segregate phenomena produced by the generator operation itself and those from the surrounding stray fields noise. Fig. 2 shows the 2-D machine calculation domain with the flux distribution at no-load during healthy operation. 0018-9464 \u00a9 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. The short-circuited turns in the generator rotor field winding can be experimentally imposed by applying the dc voltage only partially on the winding, as shown in Fig. 3(a). For instance, if we wish to simulate a short circuit between the points J and J2, the dc voltage is applied only between the points J2 and K"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000011_carpi.2012.6473371-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000011_carpi.2012.6473371-Figure10-1.png",
        "caption": "Fig. 10. Sequential payload lifting scenario.",
        "texts": [
            " Once a 3D model of a structure exists, the location of damages and repairs may be tracked for future servicing periods using a servicing history. Such models would allow better planning for current or future maintenance sessions and thus increase the efficiency of the servicing process in general. RWUAVs have limited payload capabilities and thus may not be suitable to perform tasks where heavier machinery is required. This includes in particular cleaning applications e.g. cleaning boiler walls using water jet technology. Nonetheless, RWUAVs may be of help for such tasks as well. Figure 10 visualizes the basic idea of a RWUAV that attaches a lightweight cable winch like system to help pulling a heavier machine to a remote location which might be inaccessible otherwise. This type of sequential payload lifting may also be useful to secure heavy climbing robots such as [8, 9]. C. Inspection by Contact For particular components pure visual health assessment may not be sufficient. In such cases NDT methods which provide additional information about material defects are employed. The most powerful and widespread technology to collect this type of information is UT (Figure 11)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001934_1.4030344-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001934_1.4030344-Figure3-1.png",
        "caption": "Fig. 3 Full ring magnetic bearing: (a) front view and (b) sectional side view",
        "texts": [
            " Due to repulsion between stator and rotor magnets, shaft remains levitated even under loaded Contributed by the Tribology Division of ASME for publication in the JOURNAL OF TRIBOLOGY. Manuscript received December 8, 2014; final manuscript received March 28, 2015; published online May 25, 2015. Assoc. Editor: Bugra Ertas. Journal of Tribology OCTOBER 2015, Vol. 137 / 042201-1Copyright VC 2015 by ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use conditions. In the present heading, the vertical forces for three configurations are: (1) full ring rotor and stator magnets as shown in Fig. 3, (2) stator with cuboidal magnets arranged as axially polarized and axially polarized full ring rotor magnet (shown in Fig. 4), and (3) RMD configuration (shown in Fig. 8) has been derived using 3D Coulombian model. Mathematical Modeling for Full Ring Rotor and Stator Axially Polarized Magnet. An assembly of axially polarized stator and rotor ring magnets is shown in Fig. 3. According to Coulombian model, axially polarized magnet is represented by two charged planes located on the front and back side as shown in Fig. 3(a) or left and right side surfaces of the ring, as shown in Fig. 3(b). The charge densities in the front plane of ring permanent magnet are \u00fer and in the backside are r. The detailed derivation of the vertical force using Coulombian approach for an axially polarized bearing (Fy,z) has been provided by Tan et al. [11] and Hirani and Samanta [12] and is represented as Fy;z \u00bc r1r2 4pl0 R za\u00f0 \u00de \u00fe R za \u00fe H B\u00f0 \u00de \u00fe R za \u00fe H\u00f0 \u00de \u00fe R za B\u00f0 \u00de\u00f0 \u00de (1) where r1 and r2 are the charge distributions of the stator and rotor magnets, respectively; and R(a) is given by R a\u00f0 \u00de \u00bc \u00f0h4 h3 \u00f0h2 h1 \u00f0R4 R3 \u00f0R2 R1 e\u00fe r12 cos h\u00f0 \u00de r34 cos h0\u00f0 \u00de\u00f0 \u00der12r34 r2 12 \u00fe r2 34 \u00fe e2 2r12r34 cos\u00f0h h0\u00de \u00fe 2e r12 cos h\u00f0 \u00de r34 cos h0\u00f0 \u00de\u00f0 \u00de \u00fe \u00f0a\u00de2 1:5 dr12dr34dhdh0 (2) where \u201ce\u201d is the eccentricity between the rotor and stator magnets, \u201cR1\u201d and \u201cR2\u201d are the inner and outer radii of the rotor magnet, and R3 and R4 are inner and outer radii of the stator magnet as shown in Fig. 3(a). H and B are the axial length of the stator and rotor magnets, respectively. za is the axial offset between the rotor and stator magnet as shown in Fig. 3(b). It can be observed that Eq. (2) has four integrations and in the present work these four integrations have been solved using trapezoidal numerical integration technique in MATLAB software. Mathematical Modeling of Stator With Cuboidal Magnets and Full Ring Rotor Magnet. Full ring rotor (inner radius R1 and outer radius R2) magnet and \u201cn\u201d number of square magnets of side \u201cL\u201d are shown in Fig. 4. 042201-2 / Vol. 137, OCTOBER 2015 Transactions of the ASME Downloaded From: http://tribology.asmedigitalcollection"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001072_s00170-017-0463-2-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001072_s00170-017-0463-2-Figure2-1.png",
        "caption": "Fig. 2 Topology modification method",
        "texts": [
            " If Zg is multiples of 6, we need to control the gear rotates a pitch angle automatically after the gear turns a circle to process next 1/Zc teeth. The essential characteristic of this design is that any point will not be influenced by the processing of any other points on the design flanks. This characteristic is quite useful for the topology modification method proposed in this paper. External tooth-skipped gear honing mentioned above is presented to solve the mid-concave profile problem. Its application could also be extended to process modified flanks. As shown in Fig. 2, the position of the contact point of the honing wheel and the workpiece could be determined by controlling the rotation of the wheel and the workpiece. The green surface and the red surface represent the unmodified flank and the modified flank, respectively. The gray surface is the position of the honing wheel processing the unmodified flank. The yellow surface is the position of the honing wheel processing the modified flank. In unmodified gear processing, the green surface and the gray surface will contact at point A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure12-1.png",
        "caption": "Figure 12: Flux of force adapted lattice structure with straightened struts in a shear-loaded beam",
        "texts": [
            " In order to avoid uparching and therefore stress peaks in the structure, the curved struts, which follow the flux of force exactly, have been replaced by straight struts. This structure has been compared to the ones presented in the previous sections. Structure Geometry. To obtain the described geometry with straight struts, the nodal points of the previous structure have been retained and the curved structure has been replaced by straight beam elements. Analogue to the investigations presented before, the diameters of the struts were set to 2 mm. The resulting structure and its constraints and loads can be seen in figure 12. The mass of this geometry is 83.58 g. Thus, the weight is similar to the curved structure and 43 % below the periodic one. Stiffness. Figure 13 shows the simulation result for the z-displacement of the structure. In order to make the results more comparable, the colour scale is the same as in figure 5 and 9. The displacement in the point of force application is 0.130 mm. Together with the applied force of 300 N, a stiffness of 2317 N/mm results for the structure. This means an improvement of 7 % compared to the curved structure and 9 % compared to the periodic geometry"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002889_s40436-016-0158-1-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002889_s40436-016-0158-1-Figure7-1.png",
        "caption": "Fig. 7 The vibration mode of the spindle unit under the static state",
        "texts": [
            " The differences of the front bearings inner ring and the outer ring thermal deformation are around 4 lm and 7 lm, respectively, and the differences of the rear bearings are around 2 lm and 4 lm. The front bearings thermal deformations are larger than the rear bearings both in radial and in axial directions. Commonly used in engineering is the first two order natural frequency of the rotor part. For the static rotor, the axial preload is 60 N, and the simulation result the first two order natural frequency is shown in Fig. 7. The first order natural frequency of the spindle unit is 804 Hz with the vibration mode swinging up and down at the right end. The second order natural frequency of the spindle unit is 1 765 Hz with the vibration mode swinging up and down in the middle. With the simulation results plugging into Eqs. (17) and (18) and the diagram provided in Ref. [16], we can get the additional load of the front or the rear bearings that is generated by thermal deformation. Then the bearing stiffness under the thermal static state can be calculated by Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002720_978-3-319-24055-8-Figure3.11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002720_978-3-319-24055-8-Figure3.11-1.png",
        "caption": "Fig. 3.11 Vibration due to mode 6 force shape, 24 cycles. An example of a new area of concern that has arisen in the transition to the detailed model, and which is being targeted through further design optimisation",
        "texts": [
            " However, it is better to design the housing with knowledge of guidance on likely noise problems rather than have to design in ignorance of where such problems may occur. 3.8 Further Simulation: Refinement The final stage of using this simulation was in the refinement of the design. The vibration from all major sources, including harmonics, was be inspected and compared across the speed range. Not all vibration results improved between the intermediate and detailed models. New peaks appeared and it was possible to inspect why these occurred. Figure 3.11 highlights an area of the structure that is particularly active in response to Mode 6 Force Shape, 24 cycles. The main area of activity is where the oil tank is to be located for the dry sump system, an area that did not exist in the intermediate model. This issue was highlighted to the design team who were able to consider mitigation steps in this particular area for the final design. Up to this point, the system response was calculated due to representative unit excitations, irrespective of speed and load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000599_j.engfailanal.2012.02.008-Figure17-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000599_j.engfailanal.2012.02.008-Figure17-1.png",
        "caption": "Fig. 17. Computational model of casting pedestal traverse with reinforcement.",
        "texts": [
            " 16. With respect to supposed areas of material degradation and dimensional possibilities on the upper side of traverse beam it was proposed replacement of parts of upper flanges in the shape of circular sections with radius 200 mm according to Fig. 16. In Fig. 16 is given photograph of upper side of traverse beam after replacement of material in the corner of opening for bearing. In order to decrease stress levels in locations K and L it was recommended to reinforce traverse beams according to Fig. 17, where is given computational model for the finite element method. In Fig. 18 is shown the field of equivalent stresses on reinforced traverse beam for the loading by one full ladle. From comparison of equivalent stress fields on upper side of traverse beam before (Fig. 15) and after (Fig. 18) reinforcement results that reinforcement decreases in locations of stress concentrators K and L stress peaks by approximately 15%. On the basis of analysis obtained during solution problems of loading of casting pedestal the following conclusions can be stated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003537_tmag.2018.2854666-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003537_tmag.2018.2854666-Figure7-1.png",
        "caption": "Fig. 7. Meshed stator and rotor with unmeshed air gap.",
        "texts": [
            " The Crank\u2013Nicolson method is a second-order method, which is implicit in time and numerically stable. The time discretization of (58)\u2013(61) in the matrix form can be written as (62) shown at the bottom of this page, where {S} = \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 {0}n\u00d71 {Us}m\u00d71 {0}26\u00d71 {0}26\u00d71 \u23ab\u23aa\u23aa\u23ac \u23aa\u23aa\u23ad and {D} = \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 {A}n\u00d71 {Is}m\u00d71 {Ir }26\u00d71 {Ur }26\u00d71 \u23ab\u23aa\u23aa\u23ac \u23aa\u23aa\u23ad . The boundary condition has been applied by forcing the magnetic vector potential (A) to be zero at the outer boundary of the machine. To incorporate the rotor movement and the rotor eccentricity, the air-gap stitching method has been applied. Fig. 7 shows the meshed region of the stator and the rotor domains. \u23a1 \u23a2\u23a2\u23a3 ([K ]t+ t + 2 t [Cr ] ) [Ps ][T ] [0] [Pr ] 2 t [Qs ] [R] + 2 t Les [0] [0] 2 t [Qr ] [0] [Rr ] [Tr ] [0] [0] [R] + 2 t Les MT 2 M2 \u23a4 \u23a5\u23a5\u23a6 {D}t+ t = \u2212 \u23a1 \u23a2\u23a2\u23a3 ([K ]t+ t \u2212 2 t [Cr ] ) [Ps ][T ] [0] Pr \u2212 2 t [Qs] [R] \u2212 2 t Les [0] [0] \u2212 2 t [Qr ] [0] [Rr ] [Tr ] [0] [0] [Rer ] \u2212 2 t [Ler ] MT 2 M2 \u23a4 \u23a5\u23a5\u23a6 {D}t + {S}t+ t + {S}t (62) It can be observed from Fig. 7 that the air-gap region has been divided into four parts. Two parts of the air gap have been modeled with the stator region and one part of the air gap has been modeled with the rotor region. To incorporate the rotation and the eccentricity condition, the remaining part of the air gap is stitched using six-nodded triangular elements, as shown in Fig. 8. Algorithm in Fig. 9 has been used for the stitching of the air-gap region. Suppose A, C, E, G, . . . are the stator side nodes and B, D, F, . . "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001723_j.mseb.2018.10.015-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001723_j.mseb.2018.10.015-Figure13-1.png",
        "caption": "Fig. 13. Schematic of the electron exchange occurring at the surface of bioelectrode in the presence of redox species.",
        "texts": [
            " The urea sensing spectra using the bioelectrode (Ur/NiO/ITO/glass) is obtained by varying the concentration from 0.83 to 16.65mM (Normal physiological range: 1.33\u20133.33mM) and is given in Fig. 12. Fig. 12 shows the rise in the value of the peak oxidation current with rise in urea concentration, this continuous rise of current indicates that more number of electrons are released during oxidation of urea by increasing urea concentration. Detailed of the sensing mechanism involved using Ur/NiO/ITO/glass as bioelectrode during the oxidation and reduction of urea is demonstrated in Fig. 13. Ammonium ion, bicarbonate ion and hydroxide ions are formed during the hydrolysis of urea immobilized by ureas. The biochemical reaction occurs over the bioelectrode is given as: + \u2192 + + + \u2212 \u2212UrsUrea 3H O 2NH HCO OH2 4 3 (2) It is important to note that urease belongs to the \u2018hydrolases\u2019 category of enzymes. The sensing mechanism of urea using Urs/NiO/ITO/ glass bioelectrode is shown in the schematic given in Fig. 13. The activity of urease is due to the SH group present on the active site of urease that can be oxidized by the chemical reaction with urea as given in Eq. (2). After oxidation, the S\u2013S bond thus produced can be reduced back to the SH group by the electrode reactions, thus recovering the original enzyme activity [31]. Hence, as shown in Fig. 13, the reduced urease loses electrons which are subsequently captured by [Fe(CN)6]3\u2212 ions present in the PBS buffer solution near the vicinity of the bioelectrode getting reduced to [Fe(CN)6]4\u2212 ions. These in turn oxidize to [Fe(CN)6]3\u2212 ions during forward CV scan at a potential of around 0.4 V and thus transferring electrons to the bottom electrode i.e. ITO in the present case via matrix utilizing the good electron communication feature of the NiO thin film matrix. During the reverse CV scan the whole process is repeated in the reverse direction as shown in Fig. 13. Thus, as the urea concentration increases, biochemical reaction takes place resulting in the release of more number of electrons and hence, a corresponding increase in CV peak oxidation current is observed. The inset of Fig. 12 shows the linear increase in value of oxidation peak current value ranges 0.87mA\u20132.00mA for urea concentration increases between 0.83mM and 16.65mM. The sensing response characteristics are repeatedly taken several times with same urea concentrations and are found to be in agreement within an accuracy of\u00b1 5%",
            "0) with Nessler\u2019s reagent (10 \u03bcl) and urea (100 \u03bcl) for different concentrations (ranges from 0.83 to 16.65mM) and the given bioelectrode was immersed for around 2min in the solution. Ammonia is produced by the hydrolysis of urea which produced a colored product by reacting with Nessler\u2019s reagent NH2Hg2I3 [32], the reaction is catalyzed by urease. The initial and final difference of the absorbance value for the bioelectrode was measured at 385 nm and the resulting values were plotted as a function of urea concentration is shown in Fig. 13. A linear increase in the enzyme activity values was observed with rise in concentration of urea up to 8.32mM (Fig. 14) with the saturating behaviour for higher values. The apparent enzyme activity, i.e., the amount of enzyme bound on the surface of the NiO thin film matrix, has been calculated using the equation aapp enz (U cm\u22122)=AV/\u03b5ts, where A is the difference in absorbance before and after incubation of the bioelectrode in the saturated solution, V is the total volume (3.17 cm3), s is the surface area of the electrode, \u03b5 is the millimolar extinction coefficient of Nessler\u2019s reagent and t is the reaction time (2min) [32]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000169_ext.12016-Figure13-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000169_ext.12016-Figure13-1.png",
        "caption": "Figure 13 3D finite element model of the forming tool with contour related coordinate system, u and v.",
        "texts": [
            " \u0307 = Zt Z \u03c8\u0307 (1) Low stiffness of the springs between drive and component prevents the material flow from strong influence of this, in the real process not existing, component drive. The material model was taken from the Marc & Mentat material library and follows Refs. 6 and 7. The model consists of about 4500 nodes and approximately the same number of elements. Because of the load transfer from 2D to 3D model, the total load case time is subdivided into load cases with automatic time step increment and load cases with fix small time steps. This simulation was performed under consideration of 100\u25e6C component temperature. The 3D model is shown in Fig. 13. It has almost 300,000 nodes and a similar number of elements. 34 Experimental Techniques 39 (2015) 28\u201336 \u00a9 2013, Society for Experimental Mechanics The tool gearing is reduced to three teeth. Material behavior is assumed as linear elastic. Boundary conditions prevent any axial, translational and rotatory motion of the tool. Nodal displacements in axial direction are set to 0 up to a certain diameter at the tool faces. Radial nodal displacements at the shaft hole are also prohibited. The same is considered for the tangential displacements at the flute for torque transmission"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002541_j.jsv.2016.01.019-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002541_j.jsv.2016.01.019-Figure2-1.png",
        "caption": "Fig. 2. Elastomer sample.",
        "texts": [
            " The incident light beam is divided into two beams, one is reflected by the chromium layer, and the other one goes through the silica layer and the interface before being reflected by the elastomer surface. The difference of optical path between the two beams generates an interferometric image where the pixel colour depends on the distance between the chromium layer and the elastomer. A camera with a resolution of 2448 2050 pixels is equipped with a varying zoom allowing a full size image from 7 mm 5.9 mm to 0.95 mm 0.8 mm. It records video at the rate of 9 frames per second. The elastomer sample is an extruded profile of 30 mm length, as shown in Fig. 2. The profile consists of a half disk of 10 mm diameter extended by a 10 mm right section. It has overall dimensions of 15 mm height and 10 mm width. Furthermore, a small cylinder of radius 0.5 mm is placed at the bottom of the main cylinder. The contact with the glass is achieved along a generating line of this small cylinder (x-axis). The sample is clamped in a rigid sample holder at its top side and along two vertical sides of height 6.8 mm. The glass disk is in contact with the sample bottom"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003388_b978-0-12-812959-3.00016-2-Figure16.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003388_b978-0-12-812959-3.00016-2-Figure16.1-1.png",
        "caption": "FIGURE 16.1",
        "texts": [
            " Thus, the chapter is organized as follows: after the introduction, the second section is concerned with the definition of the wind energy. The third section deals with the types of wind turbines. The modeling of the wind turbine is studied in depth in Section 4. Section 5 presents the principle of the wind turbine emulator. Finally, Section 6 summarizes the simulation results. A wind turbine is a device that transforms a part of the kinetic energy of the wind (fluid in motion) into available mechanical energy on a transmission shaft and then into electrical energy via a generator [7] (Fig. 16.1). In fact, wind energy is a nondegraded renewable energy, is geographically diffuse, and is in correlation with the season (electric power is much more demanded in winter, and it is often during this period that the average wind speed is high). Furthermore, it produces Energy conversion: from kinetic to electrical energy. no atmospheric discharge or radioactive waste [6,7]. However, it is random in time, and its capture remains complex and requires large masts and blades (up to 60 m for several megawatts) in clear areas to avoid turbulence [8]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure20.5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure20.5-1.png",
        "caption": "Fig. 20.5 Cylindrical model of a solid rotor with conducting slot wedges. (a) simplified scheme of the cross-section of the solid rotor with conducting slot wedges; (b) solid rotor with conducting slot wedges as a layered structure",
        "texts": [
            "2 Solid Rotor with Conducting Slot Wedges: The Layered Model 767 1 _\u03c1kl \u00bc 1 \u20ac\u03c1kl \u00fe 1 _\u03c1ck \u00bc 1 \u03c1kl tkl b\u03a0 \u00fe 1 \u03c1z tkl bk 1\u00fe j\u03c91\u03bccks b2k 12\u03c1z \u00f020:30\u00de where _\u03c1kl is the average resistivity of the rotor slot wedge region. Using expression (20.30) allows the rotor slot wedge region to represent as a homogeneous conducting layer. In magnetic regard, the rotor slot wedge region represents as magneto-anisotropic layer. The components of the magnetic permeability of this rotor layer (\u03bc\u03c6kl, \u03bcRkl) can be determined by the expressions obtained in (16.25) and (16.27). The rotor yoke region is considered an isotropic conducting layer. Now, the model of the solid rotor with conductive slot wedges obtains the form in Fig. 20.5b (taking into account the air gap layer with length equal to \u03b4/2). Below, based on the model in Fig. 20.5b, we consider the circuit loops of the eddy currents induced in a solid rotor with conducting slot wedges at the weak skin effect. According to Fig. 20.5b, the model of a solid rotor with conducting slot wedges represents a four-layer system. This model includes the air gap layer with a length equal to \u03b4/2, the slot wedge, the wound part of the rotor tooth and the rotor yoke layers. To obtain the equivalent circuit for a solid rotor with conducting slot wedges, it is necessary to have the equivalent circuits representing the corresponding layers of the rotor model shown in Fig. 20.5b. The rotor model layers in Fig. 20.5b can be replaced by T- or L-circuits. The air gap layer in Fig. 20.5b is replaced by the circuit in Fig. 5.5. By analogy with the T-circuit in Fig. 18.6, an equivalent circuit of the slot wedge layer acquires the form in Fig. 20.6. Based on expressions (18.24) and (18.25), the impedances of the equivalent circuit in Fig. 20.6 can be written as Zkl0 \u00bc j\u03c91\u03bc\u03c6k1Rkl1 nkl \u03beLkl \u03be\u03c4kl akl2 bkl2 \u03a8 klakl2bkl1 \u03a9klakl1bkl2 \u00f020:31\u00de Ztkl1 Ztkl2 Hkl2Hkl1 Zkl0 Ekl2Ekl1 Fig. 20.6 T-circuit representation of the solid rotor slot wedge region 768 20 Solid Rotor with Conducting Slot Wedges: Circuit Loops",
            " The rotor yoke layer is replaced by a two-terminal network, the impedance of which is calculated by the expression given in (16.32). Now, connecting in cascade the equivalent circuits representing the air gap (Fig. 5.5), slot wedge (Fig. 20.6), wound part of the rotor tooth (Fig. 16.5) and rotor yoke layers, the equivalent circuit of the solid rotor with the conducting slot wedges takes the form in Fig. 20.7. The equivalent circuit in Fig. 20.7 was constructed using T-circuits for the rotor model layers (Fig. 20.5). An equivalent circuit can be obtained for the case when these rotor layers are replaced by L-circuits. By analogy with the equivalent circuit in Fig. 18.8, the L-circuit of the slot wedge region obtains the form in Fig. 20.8. In this equivalent circuit, we have for the values of E kl2,H kl2 and Z \u03c4kl the following * *Ztkl Hkl2Hkl1 Zkl *Ekl2Ekl1 Fig. 20.8 L-circuit representation of the solid rotor slot wedge region 20.2 Solid Rotor with Conducting Slot Wedges: The Layered Model 769 E kl2 \u00bc Ekl2ckl;H kl2 \u00bc Hkl2 ckl ; Z \u03c4kl \u00bc Z\u03c4klc 2 kl \u00f020:33\u00de In accordance with the equivalent circuit in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001808_1.4037667-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001808_1.4037667-Figure1-1.png",
        "caption": "Fig. 1 Rotor-bearing configuration and the coordinate system",
        "texts": [
            " The other assumptions are as follows: (1) two bearings are concentric and (2) entire rotor is a rigid body with four degrees-of-freedom (4DOF) motions. Orbit simulations are carried out by timeintegration of the equations of the rotor 4DOF motions, equation for bump foil deflection, and transient Reynolds equation for each bearing simultaneously. Similar approach using 4DOF rigid rotor model was used in the other previous studies [24\u201326]. The 4DOF motions of the rotor are translation motions along the X and Y directions, and rotational motions about the X and Y axes (w and n, respectively). Figure 1 depicts the rotor-bearing system configuration and the global coordinate system. The equations of motions for the rotor are mR \u20acX \u00bc Fbrg X \u00fe Fimb X \u00fe mRg mR \u20acY \u00bc Fbrg Y \u00fe Fimb Y IT \u20acn \u00fe IPx _w \u00bc Mbrg n \u00feMimb n IT \u20acw IPx _n \u00bc Mbrg w \u00feMimb w (1) where mR is the rotor mass, Ip ; IT are the polar and transverse moments of inertia of the rotor, respectively. Fbrg X=Y ; Fimb X=Y are the bearing reaction forces and the imbalance forces, respectively, and Mbrg w=n ; Mimb w=n are the reaction moments from the bearings and external moments from imbalance forces, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002135_0959651813515206-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002135_0959651813515206-Figure1-1.png",
        "caption": "Figure 1. Submersible with bow and stern hydroplanes.",
        "texts": [
            " In fact, it has been shown that very large adaptation gains reduce the stability margins of the traditional adaptive systems.31 The organization of the article is as follows: Section \u2018\u2018Submarine model\u2019\u2019 presents the mathematical model and the control problem. The L1 adaptive control system is derived in section \u2018\u2018L1 adaptive autopilot design.\u2019\u2019 Section \u2018\u2018Simulation results\u2019\u2019 presents the simulation results, and conclusions are provided in section \u2018\u2018Conclusion.\u2019\u2019 For this study, the dive-plane nonlinear time-varying dynamics of submarines (as shown in Figure 1) are considered. This model is similar to the neutrally buoyant model of the submarine investigated earlier,26\u201328 but it includes hydrodynamic nonlinearities as well for a practical representation. Of course, the design approach considered here is applicable to submarines that are not necessarily neutrally buoyant. The equations of motion, describing the depth variable along the body-fixed zaxis and the pitch angle along the y-axis, are given by _w(t)= Z9wU Lm93 w(t)+ 1 m93 (Z9 _u +m9)U _u(t)+ Z9 _QL m93 _Q(t) + Z9dBU 2 m93L dB(t)+ Z9dSU 2 m93L dS(t)+ Zd(t) 0:5rL3m93 +Zn(w, q) \u00f01\u00de and _Q(t)= M9 _w LI92 _w(t)+ M9wU L2I92 w(t)+ M9 _uU LI92 _u(t) + M9dBU 2 L2I92 dB(t)+ M9dSU 2 L2I92 dS(t) + 2mg(zG zB) rL5I92 u(t)+ Md(t) 0:5rL5I92 +Mn(w, q) \u00f02\u00de where w is the velocity along the z-axis, h is the depth of the vehicle, u is the pitch angle, Q= _u is the pitch rate, and dB and dS are the hydroplane deflections in the bow and stern planes, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002749_j.ijnonlinmec.2016.08.007-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002749_j.ijnonlinmec.2016.08.007-Figure9-1.png",
        "caption": "Fig. 9. Mechanical model of the woodpecker toy.",
        "texts": [
            " Although the maximum value for \u03a9(t) is much lower than the limit given by the non-slipping condition presented above, derived for constant \u03a9, the ball slides because the friction is not sufficient to keep up with the increased variation in angular velocity, since the above condition is derived under assumptions that are not valid for this case. A common characteristic between these cases is that the ball slipping results in a spiraling trajectory (Fig. 8b and e), regardless of whether the angular speed \u03a9 is constant or periodic. The woodpecker toy, shown in Fig. 9, consists of a pole fixed to the ground, a sleeve, a torsional spring and a woodpecker. The sleeve and the woodpecker are connected with a revolute joint and a torsional spring. The hole in the sleeve is slightly larger than the width of the pole. The toy's response is heavily affected by friction and contact events. The woodpecker moves down due to gravity with a pitching motion and impacts the pole, while the Please cite this article as: A. Pournaras, et al., Dynamics of mechanica detection algorithm, International Journal of Non-Linear Mechanics ( sleeve acts as a jamming mechanism in the presence of friction",
            " The sleeve's lateral movement was assumed to be small and it was neglected. Another major assumption was that both the woodpecker and the sleeve rotate with a small angle, so that the corresponding gap functions were calculated with linearized kinematics. In the present study, the model is created by using basic solids. Specifically, the pole is represented by an orthogonal parallelepiped, the sleeve by two orthogonal parallelepipeds, the body of the woodpecker by an ellipsoid, its head by a sphere, its beak by a conical frustum and a sphere at the tip (see Fig. 9). There exist also a revolute joint and a torsional spring between the sleeve and the woodpecker. Contact is taken into account between the beak's tip and the pole (sphere to box) and between the sleeve and the pole (box to box). Values for the geometry, masses, mass moments of inertia and contact parameters are taken from reference [29]. Here, the model is also planar. However, the sleeve's lateral movement is taken into account. The major difference with [29] is that the rotations of both the sleeve and the woodpecker are not assumed to be small a priori, hence the kinematics is nonlinear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003507_978-981-13-0305-0_5-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003507_978-981-13-0305-0_5-Figure4-1.png",
        "caption": "Fig. 4 Parameters related to cost",
        "texts": [],
        "surrounding_texts": [
            "The present stage describes the procedure for chemical post processing of the specimen. A chemical bath is prepared with acetone (Acetone 85% and distilled water 15%) as ABS plastic is soluble in acetone and also due to its low cost, low toxicity and good diffusion property. The specimen is now immersed for one hour. After removing it from ethylene, it is dried and weighed and its dimensions are measured. The surface roughness of the part is again measured to find out any difference in roughness."
        ]
    },
    {
        "image_filename": "designv11_13_0001298_rpj-04-2016-0055-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001298_rpj-04-2016-0055-Figure3-1.png",
        "caption": "Figure 3 \u2013 Generic geometry of support",
        "texts": [
            " Secondly, making support in a soluble material, such as PVA or PLA, facilitating its removal without accessibility constraints. By generating a geometry incorporating a support with the minimum slope angle \u03b1, it is also possible to reduce the amount of material and therefore the manufacturing time. A tree geometry will continuously cover the entire surface to support while eliminating defects of MeshMixer software. However, it is not possible to erase the staircase effect. Therefore, the method is based on a generic shape which (Figure 3) consists of a vertical column surmounted by an inverted pyramid. This geometry will be named \u201ccolumn\u201d. In order to integrate the support optimization in the design stage, a new method dedicated to support generation (Figure 4) is proposed. It is based on an integrated tool in a CAD software. The aim of this method is to optimize the volume of support and its impact on the surface finish. In this method, steps from the discretization of the CAD model to modeling of the support do not have a physical representation in the digital mockup"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure7-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure7-1.png",
        "caption": "Figure 7: Torsion-loaded shaft composed of a helix-shaped lattice structure, based on [12]",
        "texts": [
            " In total, it can be stated that inside the horizontal struts, a large part of the overall stress is affected by bending loads. Since this is a very disadvantageous stress condition for lightweight design, the build-up of the lattice structure has to be adapted in a way, that predominantly push and pull forces appear inside the single struts. An approach to reach this state has been presented in [12]. Therein, examinations on a torsionloaded hollow shaft, which was build up out of a helix-shaped lattice structure, have been executed (see figure 7). Among others, the angles of the struts (supporting struts with same rotating direction and stiffening struts with contrary rotating direction like the torsional moment) compared to the middle axis of the shaft have been varied. Thereby, it has shown that the structure had the best stiffness to mass ratio, when the angles in both directions were 45\u00b0 compared to the rotational axis. This was explained with the help of the main stress tensors (respectively the flux of force) of a torsion-loaded hollow shaft, which also run along a helical path with an angle of 45\u00b0 compared to the shaft axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000291_amr.907.75-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000291_amr.907.75-Figure8-1.png",
        "caption": "Figure 8: Flux of force adapted lattice structure in a shear-loaded beam",
        "texts": [
            " For this, a MATLAB-based algorithm has been developed, which determines the courses of the flux of force along the first, second or third main stress direction from a given starting point. The input for this algorithm is the solver-output data of a Finite-Elements-Analysis of the design space under the defined constraints and loads. The procedure of the algorithm is based on an approach, presented in [10] for the optimization of fibre-reinforced plastics components. Structure Geometry. The course of the lattice structure inside the design space (see figure 3) has been adapted to the internal flux of force for the given constraints and loads (see figure 8). In order to keep the results comparable, the struts\u2019 diameters have been chosen to 2 mm as in case of the periodic structure. So, the course of the structure is the only differing variable in contrast to the periodic structure. The resulting structure has a mass of 83.57 g, which is 43 % less than in case of the periodic structure. Stiffness. Figure 9 shows the simulation result for the z-displacement of the structure. In order to make the results more comparable, the colour scale is the same as in figure 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002608_1464419316636968-Figure9-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002608_1464419316636968-Figure9-1.png",
        "caption": "Figure 9. Left-handed coordinate system at right rail.27",
        "texts": [
            "comDownloaded from where LTij \u00bc F Yijf2a cos\u00f0 LF \u00fe wij\u00de cos\u00f0 R wij\u00de RLF sin\u00f0 LF \u00fe wij\u00de cos\u00f0 R wij\u00de RR cos\u00f0 LF \u00fe wij\u00de sin\u00f0 R wij\u00deg \u00fe F Zijfsin\u00f0 LF \u00fe wij\u00de\u00bda cos\u00f0 R wij\u00de RR sin\u00f0 R wij\u00de g \u00feM ij sin\u00f0 LF \u00fe wij\u00de cos\u00f0 R wij\u00de \u00f017\u00de LFij \u00bc F Yijf 2a cos\u00f0 LT \u00fe wij\u00de cos\u00f0 R wij\u00de \u00fe RLT sin\u00f0 LT \u00fe wij\u00de cos\u00f0 R wij\u00de \u00fe RR cos\u00f0 LT \u00fe wij\u00de sin\u00f0 R wij\u00deg F Zijfsin\u00f0 LT \u00fe wij\u00de\u00bda cos\u00f0 R wij\u00de RR sin\u00f0 R wij\u00de g M ij sin\u00f0 LT \u00fe wij\u00de cos\u00f0 R wij\u00de \u00f018\u00de 2Rij \u00bc F YijfRLF cos\u00f0 LT \u00fe wij\u00de sin\u00f0 LF \u00fe wij\u00de RLT sin\u00f0 LT \u00fe wij\u00de cos\u00f0 LF \u00fe wij\u00deg M ijfcos\u00f0 LT \u00fe wij\u00de sin\u00f0 LF \u00fe wij\u00de sin\u00f0 LT \u00fe wij\u00de sin\u00f0 LF \u00fe wij\u00de \u00bda cos\u00f0 R wij\u00de RR sin\u00f0 R wij\u00de g \u00feM ij sin\u00f0 LF \u00fe wij\u00de cos\u00f0 R wij\u00de \u00f019\u00de 2 \u00bc \u00bd2a cos\u00f0 R wij\u00de RR sin\u00f0 R wij\u00de fcos\u00f0 LT \u00fe wij\u00de sin\u00f0 LF \u00fe wij\u00de sin\u00f0 LT \u00fe wij\u00de cos\u00f0 LF \u00fe wij\u00deg \u00fe \u00f0RLF RLT\u00de sin\u00f0 LT \u00fe wij\u00de sin\u00f0 LF \u00fe wij\u00de cos\u00f0 R wij\u00de \u00f020\u00de Also, F Y and F Z are the equivalent lateral forces and M is an equivalent roll moment given by the following expressions F Yij \u00bc F n LTyij Fn LFyij \u00f021\u00de F Zij \u00bc F n LTzij Fn LFzij Fn Rzij FSUSPZwij \u00femwg \u00f022\u00de M ij \u00bc a\u00f0F n LTzij \u00fe Fn LFzij Fn Rzij\u00de RLT\u00f0F n LTyij wijF n LTxij\u00de RRF\u00f0F n LFyij wijF n LFxij\u00de RR\u00f0F n Ryij wijF n Rxij\u00de wij\u00f0M n LTyij \u00feMn LFyij \u00feMn Ryij\u00de Iwy _ wij _ wij \u00f023\u00de The normal forces on the left and the right wheels, NLT, NLF, and NR act perpendicular to the contact patch plane and can be resolved into lateral and vertical components in the track plane. The resolved normal force components are NLTyij \u00bc NLTij sin\u00f0 LT \u00fe wij\u00de \u00f024\u00de NLTzij \u00bc NLTij cos\u00f0 LT \u00fe wij\u00de \u00f025\u00de NLFyij \u00bc NLFij sin\u00f0 LF \u00fe wij\u00de \u00f026\u00de NLFzij \u00bc NLFij cos\u00f0 LF \u00fe wij\u00de \u00f027\u00de NRyij \u00bc NRij sin\u00f0 R wij\u00de \u00f028\u00de NRzij \u00bc NRij cos\u00f0 R wij\u00de \u00f029\u00de When the wheel is making double point contact at the right rail, by using a left-handed triad coordinate system (see Figure 9), the following modifications have to be made to all equations in \u2018Double point normal forces and moments\u2019 section i\u03020 \u00bc i\u0302 ; j\u03020 \u00bc j\u0302 ; k\u03020 \u00bc k\u0302 ; 0 \u00bc ; 0 \u00bc ; 0 \u00bc \u00f030\u00de The dynamic analysis of the railway vehicle was carried out after mathematical modelling. As was mentioned before, introducing the nonlinearity of the wheel profile into equations of motion of Cheng et al.20 was essential for studying the effect of hollowing on the dynamic behaviour of the vehicle. To do so, the S1002 standard profile and the worn profiles located between 20 to 100mm from back of the flange were fitted by some polynomial functions, because this part of wheel surface is in contact with the surface of rail"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000924_978-3-319-10109-5-Figure12.1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000924_978-3-319-10109-5-Figure12.1-1.png",
        "caption": "Fig. 12.1 Conditional distribution of the leakage flux and flux values of self-induction and mutual induction of an asynchronous machine with a double-cage rotor (a) leakage flux and flux values of self-induction and mutual induction of the stator winding; (b) leakage flux and flux values of selfinduction and mutual induction of the starting winding; (c) leakage flux and flux values of selfinduction and mutual induction of the working winding",
        "texts": [
            "1), all the magnitudes are expressed in the specific system of units. Here, index A of the resistance values, voltage, currents, and emfs is omitted for convenience. In (12.1), the magnitudes of the voltage, emfs, and currents are effective values. In order to obtain the equations E1p,E2p and E4p for the emfs, the stator and rotor windings can be represented as the three inductively coupled circuits. The conditional pictures of the magnetic fields created by the currents of these circuits (I1, I2 and I4) have been presented in Fig. 12.1. These field pictures (Fig. 12.1) are thought to be caused by the currents of each individual winding. On the basis of Fig. 12.1 and the system of equations (2.87), we have for the emfs of E1p,E2p, and E4p 444 12 Double-Cage Rotor Circuit Loops: Weak Skin Effect E1p \u00bc jx1I1 \u00fe jx21I2 \u00fe jx41I4 E2p \u00bc jx2I2 \u00fe jx12I1 \u00fe jx 0 42I4 E4p \u00bc jx4I4 \u00fe jx14I1 \u00fe jx 0 24I2 \u00f012:2\u00de where x1, x2, x4 are the total reactance values of self-induction of the stator and rotor windings; x12, x21, x14, x41 are the reactance values of mutual induction for the corresponding stator and rotor windings; x 0 24, x 0 42 are the total reactance values of mutual induction of the corresponding pairs of the rotor windings. According to Fig. 12.1, the total reactance values of self-induction x1, x2, and x4 can be represented as the sum of the reactance values of self-induction and leakage reactance values of the stator and rotor windings x1 \u00bc x11 \u00fe x1\u03c3 x2 \u00bc x22 \u00fe x2\u03c3 x4 \u00bc x44 \u00fe x4\u03c3 \u00f012:3\u00de where x11, x22, and x44, and also x1\u03c3, x2\u03c3 and x4\u03c3, are the reactance values of selfinduction and leakage reactance values of the stator and rotor windings. In accordance with Fig. 12.1, the total reactance values of mutual induction of the corresponding pairs of the rotor windings x 0 42 and x 0 24 can be written as x 0 24 \u00bc x24 \u00fe x24p x 0 42 \u00bc x42 \u00fe x42p \u00f012:4\u00de where x24 and x42 are the reactance values of mutual induction of the corresponding pairs of the rotor windings; x24p and x42p are the reactance values of mutual induction conditioned by the leakage fields coupling with the corresponding pairs of the rotor windings. Taking into account (12.3) and (12.4), we can reduce the system of the equations (12",
            "7), the system of the equations (12.6) is converted as Em\u00f0 \u00de \u00bc E1m\u00f0 \u00de \u00bc E2m\u00f0 \u00de \u00bc E4m\u00f0 \u00de \u00bc jx11 I1 \u00fe I2 \u00fe I4\u00f0 \u00de \u00bc jxmIm \u00f012:8\u00de where xm\u00bc x11; Im\u00bc I1 + I2 + I4. Here, Im represents the magnetizing current and xm is the magnetizing reactance of an electric machine with a double-cage rotor. Now with the use of the expressions (12.7) and (12.8), we have the following for the system of the equations (12.5): E1p \u00bc jx1\u03c3I1 \u00fe jxmIm E2p \u00bc jx2\u03c3I2 \u00fe jx42pI4 \u00fe jxmIm E4p \u00bc jx4\u03c3I4 \u00fe jx24pI2 \u00fe jxmIm \u00f012:9\u00de Based on Fig. 12.1, the reactance values x2\u03c3, x4\u03c3, x24p, and x42p used in (12.9) take x2\u03c3 \u00bc x\u03a02\u03c3 \u00fe x\u03c4kR \u00fe x\u03c402 x4\u03c3 \u00bc x\u03a04\u03c3 \u00fe x\u03c4zR3 \u00fe x\u03c4zR2 \u00fe x\u03c4kR \u00fe x\u03c402 x24p \u00bc x42p \u00bc x\u03c4kR \u00fe x\u03c402 \u00f012:10\u00de The conditions shown in (12.8) and (12.10) have also been shared inAppendixA.12.1 for the reactance values expressed in the phase system of units. As shown in Appendix A.12.1, the reactance values used in (12.8) and (12.10) are reduced, in this case, to the stator winding. By the expressions in (12.10), the system of the equations (12",
            "2, the impedance of the working winding circuit loop takes Z 0 4 \u00bc j x\u03c402 \u00fe x\u03c4kR\u00f0 \u00dec\u03c42 \u00fe rc4 s \u00fe j x\u03a04\u03c3 \u00fe x\u03c4zR2 \u00fe x\u03c4zR3\u00f0 \u00de h i c2\u03c42 \u00f012:15\u00de We can proceed from the fact that the condition (x\u03c402 + x\u03c4kR) (rc2/s) is usually true. Taking this into account, we have from (12.14) that c\u03c42 1.0. As a result, the impedance of the working winding circuit loop arises from (12.15) Z 0 4 \u00bc rc4 s \u00fe j x\u03a04\u03c3 \u00fe x\u03c4zR3 \u00fe x\u03c4zR2 \u00fe x\u03c4kR \u00fe x\u03c402\u00f0 \u00de \u00f012:16\u00de As it follows from (12.10) and (12.15), the leakage reactance of the working winding arising from the field distribution picture created by the winding\u2019s own current (Fig. 12.1b) does not correspond, in the general case, to the leakage reactance of the working winding used in an electric machine equivalent circuit. 448 12 Double-Cage Rotor Circuit Loops: Weak Skin Effect In accordance with the expressions given in (12.10) and (12.15), and (12.16), these reactance values are equal when the condition c\u03c42 1.0 is satisfied. It follows from (12.14) that the condition c\u03c42 1.0 means the reactance values x\u03c402 and x\u03c4kR, representing, in this case, the reactance values of mutual induction caused by the rotor leakage fields (Fig. 12.1b), are relatively low in relation to the impedance of the starting winding. By the expressions given in (12.13) and (12.16), the equivalent circuit in Fig. 12.2 can be converted to the form in Fig. 12.3. This equivalent circuit represents a three-loop circuit with a parallel connection of the elements, which is the basis from which the system of the voltage equations arises U1 \u00bc r1 \u00fe jx1\u03c3\u00f0 \u00deI1 \u00fe jxmIm 0 \u00bc rc2 s \u00fe j x\u03a02\u03c3 \u00fe x\u03c4kR \u00fe x\u03c402\u00f0 \u00de h i I2 \u00fe jxmIm \u00bc rc2 s \u00fe jx2\u03c3 I2 \u00fe jxmIm 0 \u00bc rc4 s \u00fe j x\u03a04\u03c3 \u00fe x\u03c4zR3 \u00fe x\u03c4zR2 \u00fe x\u03c4kR \u00fe x\u03c402\u00f0 \u00de h i I4 \u00fe jxmIm \u00bc rc4 s \u00fe jx4\u03c3 I4 \u00fe jxmIm \u00f012:17\u00de where x2\u03c3\u00bc x\u03a02\u03c3 + x\u03c4kR + x\u03c402; x4\u03c3\u00bc x\u03a04\u03c3 + x\u03c4zR3 + x\u03c4zR2 + x\u03c4kR + x\u03c402",
            "1) can be reduced to the following form E1p \u00bc jx1\u03c3I1 \u00fe jxmIm E 0 2p \u00bc jx2\u03c3k2Ek2II 0 2 \u00fe jx42pk2Ek4II 0 4 \u00fe jxmIm E 0 4p \u00bc jx4\u03c3k4Ek4II 0 4 \u00fe jx24pk4Ek2II 0 2 \u00fe jxmIm \u00f0A:12:16\u00de In (A.12.16), we use the following values x 0 42p \u00bc x42p k4I k2I ; x 0 24p \u00bc x24p k2I k4I \u00f0A:12:17\u00de Now the equations (A.12.16) take the form E1p \u00bc jx1\u03c3I1 \u00fe jxmIm E 0 2p \u00bc jx2\u03c3k 2 2I 0 2 \u00fe jx 0 42pk 2 2I 0 4 \u00fe jxmIm E 0 4p \u00bc jx4\u03c3k 2 4I 0 4 \u00fe jx 0 24pk 2 4I 0 2 \u00fe jxmIm \u00f0A:12:18\u00de where k22 \u00bc k2Ek2I; k 2 4 \u00bc k4Ek4I. On the basis of Fig. 12.1, we have for the reactance values x2\u03c3, x4\u03c3, x 0 42p and x 0 24p used in (A.12.18) x2\u03c3 \u00bc x\u03a02\u03c3 \u00fe x\u03c4kR \u00fe x\u03c402 x4\u03c3 \u00bc x\u03a04\u03c3 \u00fe x\u03c4zR3 \u00fe x\u03c4zR2 \u00fe x\u03c4kR \u00fe x\u03c402 x 0 24p \u00bc x24p k2I=k4I\u00f0 \u00de \u00bc x\u03c4kR \u00fe x\u03c402 x 0 42p \u00bc x42p k4I=k2I\u00f0 \u00de \u00bc x\u03c4kR \u00fe x\u03c402 \u00f0A:12:19\u00de We determine the reactance values x2\u03c3, x4\u03c3, x 0 24p and x 0 42p shown in (A.12.19). As it follows from (A.12.19), the reactance x2\u03c3 is determined by a permeance factor equal to \u03bb2\u00bc \u03bb 0 \u03a02 + \u03bb 0 \u03c4kR + \u03bb\u03c402. The value of \u03bb2 can be used in the expression (A.12.4) for the reactance x22"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001294_tpel.2017.2782804-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001294_tpel.2017.2782804-Figure1-1.png",
        "caption": "Fig. 1. Six-phase IPMSM: (a) coil layout and (b) direction of current vectors: Rotor angle \u03b8 refers the angle between the direct axis of the rotor and the magnetic axis of phase a.",
        "texts": [
            " Further, \u03bb\u03b1\u03b2 is transformed the synchronous dq frame via F(\u03b8) =  cos \u03b8 sin \u03b8 \u2212 sin \u03b8 cos \u03b8  , (9) such that \u03bbedq = LM \u2212 3 2L\u03b4 0 0 LM + 3 2L\u03b4  iedq + \u03c8pm 1 0  , (10) where \u03bbeTdq = F(\u03b8)[\u03bb\u03b1, \u03bb\u03b2 ]T and ieTdq = F(\u03b8)[i\u03b1, i\u03b2 ]T . These are well established equations and were appeared in textbooks, for example in [25], [26]. The six-phase motors considered in this paper have two sets of three phase windings which are separated by \u03c0/6 in electrical angle. The first group is named abc and the second group is xyz. Fig. 1 shows schematic diagram of two pole six-phase IPMSM. The same modeling approach is applied here as in the three phase case. In the following developments, we look at the windings of six phase IPMSM as two sets of three phase windings. The reason is to take the advantage of 0885-8993 (c) 2017 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. This article has been accepted for publication in a future issue of this journal, but has not been fully edited"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000704_sas.2012.6166307-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000704_sas.2012.6166307-Figure2-1.png",
        "caption": "Fig. 2. The probe spatial reception lobe is assumed to be focused on a small, limited portion of the cogwheel profile.",
        "texts": [
            " (2) Moreover, from an electrical point of view, the probe is modeled as a linear time-invariant system with a generic bandpass transfer function H(\u03c9) = jG\u03c9 (j\u03c9+\u03c9L)(j\u03c9+\u03c9H ) . The therm G is a gain factor taking into account the probe distance from the cogwheel, whereas \u03c9L \u03c9H are the two pulsations associated with the high-pass and low-pass cut-off frequencies. An analytical description of the probe voltage signal Vx can be obtained following the approach discussed in this Section. Referring to Fig. 2, we assume the probe spatial reception lobe to be focused on a small, limited portion of the cogwheel profile, i.e, \u03b10 \u03b80 (the assumption typically makes sense, since D R). By denoting with r : [0, 2\u03c0) \u2192 R + the polar representation of the cogwheel profile (with respect to the point C), we note that the cogwheel profile can be locally approximated by the projection ry : R \u2192 R + of the polar 978-1-4577-1725-3/12/$26.00 \u00a92012 IEEE profile r on the y axis, in front of the probe. Accordingly, by defining ry(y) = r ( y R mod 2\u03c0 ) , (3) we obtain an unrolled profile ry that is periodic with period 2\u03c0R"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000599_j.engfailanal.2012.02.008-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000599_j.engfailanal.2012.02.008-Figure1-1.png",
        "caption": "Fig. 1. Casting pedestal with ladles.",
        "texts": [
            " From the results of strain-gage measurements accomplished under simulated and operational regimes can be concluded that supporting structure of pedestal is exposed to considerable dynamical loading. On the basis of analysis was found out that taking into account previous loading history of pedestal, the lifetime of traverse beam is exhausted. Consequently, the modifications of pedestal structure were proposed with the aim to allow its further operation. 2012 Elsevier Ltd. All rights reserved. Casting pedestal (Fig. 1) is one of the most important equipment of steelworks with continuous casting of slabs. In the technological flow it serves for transportation of ladles with liquid steel between converters and tundish of machine for continuous casting. During the operation, the structure of casting pedestal is exposed to dynamic loading due to transported mass of ladles with liquid steel, because the pedestal carries out on roller bed reversible rotational movement around vertical axis ov (by 180 ) and the traverse of pedestal tilts around horizontal axis oh (by approximately 7.0 ) (Fig. 1). The welded carrying system of casting pedestal consists of middle part and two beams (Fig. 2). The traverse beams are connected with middle part by bolt and wedge joints. After more than 10 years of operation of casting pedestal arose the problems with releasing of bolt and wedge joints. Subsequently, there were found out cracks in the supporting structure of pedestal that lead to proposals of modifications of the structure. In the paper are given results obtained during assessment of crack initiation causes in the supporting structure of pedestal as well as the measures for ensuring its further operation",
            " Ladle in hinges C, D (mostly with rest of steel) is moving to reserve position from which is taken (hinges C, D are without ladle). Short time before end of steel casting from ladle in hinges A, B full ladle is given to hinges C, D and the process repeats. During operation the traverse of casting pedestal provides rickety rotational movement still by 180 around axis ov. During the casting traverse provides additional rickety rotational movement around horizontal axis oh so that ladles are moved in vertical direction. Rotational movement around vertical axis ov is realized by rollers on which whole casting pedestal (Fig. 1) lies, rotational movement around horizontal axis oh is realized by teeth on the arms with toothing and four hydraulic pistons [4]. Beams of traverse are jointed with the middle part of traverse by bolts and sunk keys (Fig. 4) that are during operation of casting pedestal (as a result of alternating and irregular loading of its arms) released. The supporting elements of traverse are composed of sheets joined by welds. Beams of traverse have closed (box) crosssection with thickness of walls that varies from 16 to 22 mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003627_icma.2018.8484453-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003627_icma.2018.8484453-Figure4-1.png",
        "caption": "Fig. 4 Curve line-of-sight guiding angle",
        "texts": [],
        "surrounding_texts": [
            "A. Single UUV Kinematics Path Following Control Based on the line-of-sight guidance[11] method and the path-tracking error model, path tracking controller is designed for each UUV.Curve line of sight guide angle is as follows: Center the point P on the path Q to create a moving coordinate system Serret-Frenet , (w,0) is the position of the next target in the moving coordinates. w is 3 times the length of UUV. The line of sight angle can be defined as: arctan( )los los y y w \u03d5 \u2212 = . (21) The first step: line-of-sight guidance The line-of-sight guiding angle los\u03d5 is e\u03d5 .a extrution function can be used to approximate the line-of-sight guiding angle. k \u03b4 is a gain factor. 2 2 1( ) 1 e e k y e a k y e y e \u03b4 \u03b4 \u03b4 \u03c8 \u2212= \u2212 + . (22) on the expected path Q ,the kinematic control law T k sruU ],,[= is deduced by the backstepping techniques.The line-of-sight guiding angle is taken as a physical inverse sine function . 2arcsin( ), ,e e los e e k y k R y \u03b4 \u03d5 \u03b5 \u03b5 +\u2212 = = \u2208 + . (23) e\u03d5 is the actual guiding angle. The Lyapunov function can be selected as follows: 21 ( ) 2nav eV \u03d5 \u03b4= \u2212 . (24) the derivative of the above formula can be obtained as: ( )( ( ) )nav eV r c s s\u03d5 \u03b4 \u03b2 \u03b4= \u2212 + \u2212 \u2212 . (25) therefore, the navigation control rate can be defined as: 1( ) ( )er c s s k\u03b4 \u03b2 \u03d5 \u03b4= \u2212 + \u2212 \u2212 . (26) Step 2: Path Tracking Position Error Combined with (11), the quadratic Lyapunov function is as follows: F e FT epos nnV = . (27) the derivative of the above formula can be obtained as: cos coscos sin 0 ( , ) sin cos sin sin cos sin t e M FF FFT F pos e e e e F F t e M F t e e t e e M e v v V n n x y v v v x v y v x \u03d5 \u03d5\u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u03d5 \u2212 = = + \u2212 \u2212 = + \u2212 . (28) the movement speed Mv of the path reference point is input as an extra degree of freedom, and the control law is selected as: 2cosM t e ev s v k x\u03d5= = + . (29) the derivation of the Lyapunov function of the path tracking error, combining the control law (29), can be obtained as: 2 2sinpos t e eV v y k x\u03b4= \u2212 . (30) under the control law of the first-step designed angular velocity, the navigation angle \u03b4 defined by (23) can be obtained as: 0 2arcsin( ), ,e e e e k y k R y \u03b4 \u03d5 \u03b5 \u03b5 +\u2212 = = \u2208 + . (31) where 0k is a proportional coefficient that adjusts the navigation angle change trend. The derivation of the path-tracking error of the Lyapunov function can further be written as: 2 2 2 2 0 22sin 0e pos t e e t e e y V v y k x k v k x y \u03b4 \u03b5 = \u2212 =\u2212 \u2212 \u2264 + . (32) where ( , ) (0,0)e ex y = is the balance point of the system, and the path of system will gradually converge on the invariant set: { }2| ( , ) 0 ,pos e e ex y \u03d5 \u03b4\u03a9 = = . (33) Step 3: kinematics path following control law The system has convergence properties under the navigation control law (26) and the tracking error control law (29).If the UUV generalized position vector starts from any point within 3{ }R\u03a9 = .Under the influence of navigation control law (26), the path gradually approaches the invariant set { }2| ( , ) ,nav x y R \u03d5 \u03b4\u03a9 \u2208 = .In the invariant nav\u03a9 and the tracking error control law (29), path tracking error also gradually converge to the invariant set { }2| ( , ) ,nav x y R \u03d5 \u03b4\u03a9 \u2208 = . combination with (26) and (29), the control law[12] that satisfies the kinematics path tracking requirement can be written as: 2 2 cos ( ) ( ) M u d k v t e e r e u U s v k x r c s s k \u03b1 \u03b1 \u03d5 \u03b1 \u03b4 \u03b2 \u03d5 \u03b4 = = = = + = = \u2212 + \u2212 \u2212 . (34) wher 1k and 2k any positive control gain, and 0tv > is controlled by the control law (34). The system path eventually converges to the balance point globally. B. Single UUV Dynamic Path Tracking Control The dynamic control law [ ]dU F, , Ts= \u0393 is deduced by a backstepping techniques.Using the backstepping techniques to expand the kinematics controller to the dynamic stage. For the path reference point Mv s\u03b1 = , its speed is a perfect tracking, only need to design the u\u03b1 r\u03b1 with backstepping techniques.this paper suppose that u , r is the virtual control input, u\u03b1 , r\u03b1 are the virtual control rate, and the error variable can be defined as: u u r r z u z z r \u03b1 \u03b1 \u2212 = = \u2212 . (35) in the basic Lyapunov function of kinematics, increasing the linear quadratic type of the velocity error variable z: 1 2 T dyn kinV V z Mz= + . (36) where ,u rm m has been given in the dynamics model. M can be defined as: 0 M 0 u r m m = . (37) the derivation of the Lyapunov function: 2 2 2 2 1 2 ( )( ) sin ( ) ( ( )) ( ) sin dyn e e t e e u u u r r r e r r r r e u u u u t e e V vy k x mz z mz z k z mr m z mu m v y k x \u03d5 \u03b4 \u03d5 \u03b4 \u03b4 \u03d5 \u03b4 \u03b1 \u03d5 \u03b4 \u03b1 \u03b4 = \u2212 \u2212 + \u2212 + + =\u2212 \u2212 + \u2212 + \u2212 + \u2212 + \u2212 . (38) taking the control law F,\u0393 as follows: F u u u r r r m d m d \u03b1 \u03b1 = + \u0393 = + . (39) where , ,u r f\u03b1\u03b1 \u03b1 can be defined as: 2 3( ) 5( ) 1 ( / )(cos ) r e r ur v f k z r k m m \u03b1 \u03d5 \u03b4\u03b1 \u03b2 \u2212 \u2212 \u2212 \u2212 = \u2212 . (40) 4( ( ))u d du k u u\u03b1 = \u2212 \u2212 . (41) 2 1 2 2 ( ) ( ) ( ) ( )t ur v e t v vt v m duvf k c s s c s s ur v m mv\u03b1 \u03b2\u03b4 \u03d5 \u03b4= \u2212 \u2212 + + + + + + . (42) where 3 4 5, ,k k k is any positive control increment. The result of the expression is: 2 2 2 2 2 1 2 3 4 0 2( ) 0e dyn e e u u r r t e y V k k x k m z k m z k v y \u03d5 \u03b4 \u03b5 =\u2212 \u2212 \u2212 \u2212 \u2212 \u2212 \u2264 + . (43) In the invariant set,the limit is taken for (38),The result is as follows: l i m 0d y nt V \u2192 \u221e = . (44) the horizontal error, longitudinal error, and velocity error of the path tracking can be converged to the invariant set. yaw error also can be converged to navigation angle. lim( , , , ) (0,0,0,0)T T e e u rt x y z z \u2192\u221e = . (45) l im li met t \u03d5 \u03b4 \u2192 \u221e \u2192 \u221e = . (46) In the invariant set,the limit is taken for (31).The result is as follows: l i m 0 t \u03b4 \u2192 \u221e = . (47) Therefore, the global path of the system gradually converges to the balance point. lim( , , , , ) (0,0,0,0,0)T T e e e u rt x y z z\u03d5 \u2192\u221e = . (48)"
        ]
    },
    {
        "image_filename": "designv11_13_0002559_tmag.2014.2327954-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002559_tmag.2014.2327954-Figure1-1.png",
        "caption": "Fig. 1. (a) Schematic and (b) simplified analytical model of AFPMC.",
        "texts": [
            " To overcome this problem, in this paper, the analytical solution is predicted based on a magnetic vector potential and the analytical model using the two polar coordinate systems corresponding to the upper PM and lower PM. The prediction involves a governing equation for the air gap (I) and the PM region (II) with axial magnetization. On the basis of a Maxwell stress tensor method, the torque of the AFPMC is computed using the solutions for magnetic flux density caused by the PMs. The analytical results are extensively validated by the 3-D FE analysis and also verified by an experimental testing of the torque produced in an AFPMC. Fig. 1(a) shows an AFPMC with axially magnetized eightpole PMs surface mounted on the iron yoke and comprising symmetric disc-type PMs of the same size. The analytical solutions are derived based on the assumptions that the permeability of the iron yoke is infinite and the relative permeability of each PM is unity [8]. Two polar coordinate systems are employed to compute the magnetic fields caused by the PMs in our analytical model, as shown in Fig. 1(b). Here, za and zb are the axial positions of the top and bottom boundaries of the upper PM. zc, zd , and ze are the axial positions of the middle of the air gap, the top and bottom boundaries of the 0018-9464 \u00a9 2014 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. lower PM. (r, \u03b8, z) and (r \u2032, \u03b8 \u2032, z\u2032) represent the upper and lower PM coordinate systems, respectively",
            " As compared with the method using one polar coordinate system, the number of analysis regions here is reduced, because the lower PM appears in the air-gap region in the coordinate frame for the upper PM, (r , \u03b8 , z). The magnetic field produced by the upper PM can thus be derived. As the two polar coordinate systems are superimposed over one another, the magnetic field produced by the lower PM can be easily obtained. On the other hand, one PM rotor of AFPMC is rotated with the other at the synchronized same rotor speed because the applied load torque is generally lower than peak torque of the AFPMC. Thus, the analytical model shown in Fig. 1(b) does not consider eddy current terms produced by relative speed between two rotor parts of the AFPMC. However, there is a possibility to occur higher load torque than peak torque of the AFPMC due to unexpected system failure. Since this model might not be capable to consider the effect of oscillations due to eddy current in the system, it is stated that [9] and [10] can be referred to solve this problem. To obtain magnetic fields produced by the upper PM, the lower PM region is considered a part of the air gap",
            " In this case, the (r , \u03b8 , z) coordinate system for the upper PM is switched to the (r \u2032, \u03b8 \u2032, z\u2032) coordinate system for lower PM, even as the upper PM is assumed to be contained in the air-gap region, as mentioned earlier. At any given position along the z-axis direction, magnetic flux density can be obtained by superimposing the two coordinate systems. Accordingly, the total magnetic flux density produced by the upper and lower PM can be derived as the algebraic sum of the magnetic flux densities obtained for each coordinate system, for both the PMs. For instance, total magnetic flux density of normal and tangential is expressed in (7) at the axial position za of the Fig. 1 Bza sz = { BII z (r, \u03b8, 2hm + hg) + BI z(r \u2032, \u03b8 \u2032, 0) } Bza s\u03b8 = { BII \u03b8 (r, \u03b8, 2hm + hg) + BI \u03b8 (r \u2032, \u03b8 \u2032, 0) } . (7) The force (F) caused by the upper PM and lower PM is induced at the bottom surface of the upper PM, at z = zb. Using the Maxwell stress tensor, the magnetic torque (T), produced by the mutual effect of both PMs, is calculated from [12] as F = \u2212S \u2329 T II z\u03b8 \u232a \u03b8 = \u2212S\u03bc0 \u2329 HII sz ( r, \u03b8, zb ) HII s\u03b8 ( r, \u03b8, zb )\u232a \u03b8 F = \u2212 S \u03bc0 [ BII sz ( r, \u03b8, zb ){ BII s\u03b8 ( r, \u03b8, zb )}\u2217] (8) T = Rm \u00d7 F"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002720_978-3-319-24055-8-Figure6.12-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002720_978-3-319-24055-8-Figure6.12-1.png",
        "caption": "Fig. 6.12 Serial design (left) in comparison to active exhaust layout (right)",
        "texts": [
            " In consequence, this means that the speaker chassis needs to be designed for a high membrane displacement. For some cars\u2014smaller gasoline and some diesel engines\u2014passive mufflers will completely be replaced by an active system in combination with some small resonators. For more powerful engines, for example 4-cylinder turbocharged gasoline engines with up to 200 kW, there will still be the need for passive muffler volume\u2014 which then can be reduced for example from 30 to 12 l. This requires that cancellation is working well. In Fig. 6.12 the comparison of a possible system layout, which is currently implemented on the internal demonstration vehicle, can be seen. The shown volume reduction of the passive system correlates to a weight saving of 4\u20135 kg, without taking the additional weight of the speaker into account. That means that in this case an overall weight reduction of nearly 1 kg could be achieved. In Fig. 6.13 the measured results obtained on the rollerbench under full-load condition is shown and compared to simulated data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003386_j.apm.2018.02.024-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003386_j.apm.2018.02.024-Figure5-1.png",
        "caption": "Fig. 5. The geometry model.",
        "texts": [
            " Given that the oil film thickness is about 1 microns, the thickness of the Eulerian body is defined as 20\u03bcm to make sure there is enough oil. When the sliding body contacts with the fixed body and slides along the x direction under a certain load, some of the oil will be squeezed out from the contact zone gradually. During this process, the sliding body pushes the excess oil ahead of it until it has moved over the entire length of the fixed body. Considering the symmetry of the geometry and boundary conditions in the z direction, only half of the structure is modelled in this study and the symmetric surfaces are labeled as C in Fig. 5. The relative positions of the three bodies are as shown in Fig. 5. ACCEPTED M ANUSCRIP T The sliding body and the fixed body are meshed with Lagrangian elements in which nodes are fixed within the material, and the oil is meshed with Eulerian elements in which nodes are fixed in space (see Fig. 6). In the contact zone, 5\u00b5m, 3\u00b5m and 1.5\u00b5m elements are used to investigate which size of element is required to adequately capture the contact behaviour. The materials of the sliding body is forged carbon steel and that of the fixed body is chilled cast iron. Due to the stress concentration, plastic deformation is most likely to occur in the real contact of the asperities",
            " ( )ref refp p \u0393 e e   , (20) 0s pU c sU  , (21) where p is the pressure, e is the internal energy, pref and eref are the reference pressure and internal energy, \u0393 is the Mie-Gr\u00fcneisen ratio and 0 0 \u0393 \u0393    in which \u03930 is a material constant and \u03c10 is the reference density, Us is the shock velocity, Up is the particle velocity, c0 is the zero-pressure ACCEPTED M ANUSCRIP T isentropic speed of sound, and s is a dimensionless parameter which is related to the pressure derivative of the isentropic bulk modulus. In this study, the parameters s and \u03930 are simplified to 0 [31]. The material property parameters of the lubricant used in this study are listed in Table 3. Table 3. Material properties of the oil. Density \u03c1lub (kg/m 3 ) Dynamic viscosity \u03b7 (Pa\u2022s) Sound velocity c0 (m/s) \u03930 s 880 0.0885 2135 0 0 The boundary conditions applied to the model includes: the z-symmetric displacement constraint is applied on the middle planes of the model which are labeled as C in Fig. 5; the bottom of the fixed body is fixed; the pressure is applied to the top of the sliding body; the sliding velocity of v=8000mm/s along the x direction is applied to the sliding body and the sliding time duration is from the moment at which the sliding body comes into contact with the fixed body to that when the two bodies separate from each other; the velocities along the normal directions of the side surface labeled as A and top and bottom surfaces labeled as B of the Eulerian body in Fig. 5 are set to zero to restrain the oil from flowing out of these surfaces; the solid\u2013solid contact is applied between the contacting rough surfaces of the sliding body and the fixed body; the fluid\u2013solid contact is applied between the sliding body and the Eulerian body and between the fixed body and the Eulerian body. Using the model built in Section 4.2, the explicit dynamic simulation is implemented. During the sliding body slides along the fixed body in the x-direction, solid\u2013solid contacts occur between some of the roughness asperities of the two bodies",
            " [32] W. Qin, Z. Zhang, Simulation of the mild wear of cams in valve trains under mixed lubrication, P. I. Mech. Eng. J - J Eng. 231 (5) (2017) 552\u2013560. ACCEPTED M ANUSCRIP T List of figure captions Fig. 1. LogS \u0336 log\u03b7 plot. Fig. 2. Flow chart of the simulation method. Fig. 3. The measured roughness heights of the contact surfaces of the specimens: (a) upper specimen; (b) lower specimen. Fig. 4. The modelled 200\u03bcm long profile heights for the upper and lower bodies: (a) upper body; (b) lower body. Fig. 5. The geometry model. Fig. 6. The mesh model. Fig. 7. Solid\u2013solid contact pressure. Fig. 8. Volume fraction of the oil in the Eulerian body. Fig. 9. Dynamic solid\u2013solid contact forces with various loads. Fig. 10. Nominal oil film thickness. Fig. 11. Change of wear coefficient with the oil film thickness. Fig. 12. Dynamic solid\u2013solid contact forces with various dynamic viscosity values of the oil. Fig. 13. Change of wear coefficient with the dynamic viscosity of the oil."
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001806_978-3-319-14418-4_177-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001806_978-3-319-14418-4_177-Figure1-1.png",
        "caption": "Fig. 1 Model of a pendulum. Symbols: M is pendulum mass, \u03b8 is angle relative to vertical, L is distance from the suspension point to the center of mass, I is moment of inertia about the suspension point, g is gravitational acceleration, \u03c4 is a torque acting on the pendulum",
        "texts": [
            " For ease of computer implementation, most investigators make use of the fact that any higher-order differential equation may be expressed as a system of simultaneous first-order equations. To provide a concrete example, consider a physical pendulum of mass M, where \u03b8 is the angle relative to vertical, L is the distance from the suspension point to the center of mass, I is the moment of inertia about the suspension point, g is the gravitational acceleration, and \u03c4 is a torque acting on the pendulum (Fig. 1). This system has the following equation of motion \u20ac\u03b8 \u00bc MgL sin \u03b8 \u00fe \u03c4\u00f0 \u00de=I (4) and we can define \u03c4 \u00bc u\u03c40 (5) such that u is the control variable, bounded between 1 and 1, that scales \u03c4 between its maximum negative (clockwise) and positive (counterclockwise) magnitude, \u03c40. If we introduce a new variable \u03c9, where \u03c9 \u00bc _\u03b8, then Eq. 4 may be rewritten as the following system of first-order differential equations _\u03c9 \u00bc MgL sin \u03b8 \u00fe \u03c4\u00f0 \u00de=I (6) _\u03b8 \u00bc \u03c9: (7) In this example, the state equations are given by Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002780_1.4034604-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002780_1.4034604-Figure2-1.png",
        "caption": "Fig. 2 Cylindrical test rolling bearings: (a) FAG N208-E-TVP2 and (b) FAG NU208-E-TVP2",
        "texts": [
            " 2 gives the description of the experimental setup and the types of rolling bearing faults diagnosed in this study. Section 3 gives the details of the feature extraction methods. Section 4 gives some insights to the data mining approaches used for fault diagnosis. Section 5 gives the results and a short discussion and Sec. 6 concludes the study. As given in Fig. 1, the experimental setup includes an AC servo motor with a 250\u20135000 rpm speed range and a load capability of 5 kN. Two split-type cylindrical rolling bearings (Fig. 2), an FAG N208-E-TVP2 and an FAG NU208-E-TVP2, were used with their original fault-free state and prescribed defects engraved onto the inner and outer races. A coupler connects the motor shaft to the bearing shaft. 2.1 Experimental Procedure. A number of experiments were done for each state of the rolling bearings including two fault-free, two inner-race defect (IRD), and two outer-race defect (ORD) states given in Table 1. The two fault-free states are combined and accepted as one fault-free state; thus, our dataset includes five final operational states"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001377_j.compeleceng.2018.02.012-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001377_j.compeleceng.2018.02.012-Figure8-1.png",
        "caption": "Fig. 8. Inclination with the vertical plane.",
        "texts": [
            " If B is above the plane of equation z = z A , i.e., z \u2019 B \u2265 0 or z B \u2265 z A , then \u03b4 = \u03b1 and \u03d5 = \u03c0 2 \u2212 \u03b1; 2. If B is below the plane of equation z = z A , i.e., z\u2019 B < 0 or z B < z A , then \u03b4 = - \u03b1 and \u03d5 = \u03c0 2 + \u03b1, with \u03b1 = si n \u22121 ( | z B \u2212z A | AB ) . 5.2.4. Calculation of the longitudinal angle \u03b8 Similarly to the zenithal angle \u03d5, the longitudinal angle \u03b8 is strongly related to the angle \u03b11 formed by the vector \u2212\u2192 AB and the vertical plane ( X\u2019, A, Z\u2019 ). Here again, we can have many cases, as summarized by Fig. 8 a. Fig. 8 b is a 2 D projection of Fig. 8 a on the plane ( X\u2019, A, Z\u2019 ). If point B is in sector 0, then \u03b8 = \u03b11 ; else if B is in sector 1, then \u03b8 = \u03b12 = \u03c0 \u2212\u03b11 ; else if B is in sector 2, then \u03b8 = \u03b13 = \u03b11 + \u03c0 ; else if B is in sector 3 then \u03b8 = \u03b14 = 2 \u03c0 \u2212\u03b11 . Therefore, we can calculate the value of \u03b8 if we know the relative position of point B and the value of inclination \u03b11 formed by \u2212\u2192 AB and ( X\u2019, A, Z\u2019 ). Calculation of \u03b11 : Let ( P 2 ): y = y A be the plane ( A, i , k ) . Let H 2 be the orthogonal projection of B on the plane ( P 2 ) ( Fig. 8 b). The distance of B to plane ( P 2 ) noted d ( B , P 2 ) is d( B, P 2 ) = B H 2 = | y B \u2212y A | 2 \u221a 0 2 + 0 2 + 1 2 = | y B \u2212 y A | \u21d2 \u03b11 = si n \u22121 ( | y B \u2212y A | AB ) . Knowing the value of \u03b11 , we can state Corollary 5 . Corollary 5. The longitudinal displacement angle \u03b8 of an actuator from point A ( x A , y A , z A ) to point B ( x B , y B , z B ) of the ( sink, i , j , k ) reference is given by the following: 1. If B is in the sector 0, i.e. , (x \u2032 B > 0 and y \u2032 B \u2265 0 ), then \u03b8 = \u03b11 2. If B is in the sector 1, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002322_j.triboint.2015.10.007-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002322_j.triboint.2015.10.007-Figure8-1.png",
        "caption": "Fig. 8. Photograph of the experimental apparatus used to measure the frictional torque and water leakage of the examined shaft seals.",
        "texts": [],
        "surrounding_texts": [
            "Three types of shaft seals (the proposed water-sealing system, an oil seal, and a mechanical seal) were evaluated using the experimental apparatus shown in Figs. 7 and 8. The test conditions are summarized in Table 2. The installation and operating conditions of the water-sealing system were those described in 2.4.1. Two commonly used seals were used for the comparison, an oil seal (AE1709A, NOK Corp., Japan) and a mechanical seal (US-1, CCU30-PAP00-303, Nippon Pillar Packing Co., Ltd., Japan), which were installed in holders, as shown in Figs. 11 and 12. The oil seal was a typical one, had an acrylonitrile butadiene rubber seal lip and was guaranteed performance above 0.03 MPa. The tested mechanical seal is commonly used for water, seawater, and oil pump components and has one compression coil spring. A maximum pressure of 1.0 MPa and circumferential velocity of 15 mm/s were guaranteed. The duration of the test was 8 h and the pressure applied by the water to the shaft seals was increased every 1 h (Fig. 13) to investigate their performance against the pressure."
        ]
    },
    {
        "image_filename": "designv11_13_0001432_s00170-018-1990-1-Figure23-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001432_s00170-018-1990-1-Figure23-1.png",
        "caption": "Fig. 23 The deformation analysis of riveting bars. a The sketch of the riveting structure. b The mesh condition of the riveting bars",
        "texts": [
            " To reduce the impact of structural deformation on riveting quality, the (a) Top surface of the sheets. (b) Faying surface of the sheets. (c) Bottom surface of the sheets. Fig. 20 Residual radial stress. a Top surface of the sheets. b Faying surface of the sheets. c Bottom surface of the sheets mapping relationship between slug rivet, aircraft panels, and automatic drill-rivet machine is built by numerical method. To ensure the analysis accuracy, full-scale model is established, because the structure of the machine is unsymmetrical. Figure 23 shows the deformation of riveting bars. To improve the accessibility, the riveting bars at the upper and lower end executors are designed as slender rods. But, the slender rod will deform under the effect of huge riveting force. The deformation will reduce the riveting quality. It is necessary to analyze the riveting bar deformation and provide an effective quality compensation approach. The structures of the riveting bars are symmetrical. To save the calculation time without reducing accuracy, symmetry model is built in the study"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002027_j.colsurfb.2014.03.047-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002027_j.colsurfb.2014.03.047-Figure2-1.png",
        "caption": "Figure 2",
        "texts": [
            "25 V appeared a 19 broad reduction peak in the cathodic scan corresponds to the electron transfer from solution to the 20 electrodeposited ploy-beryllon II film [22].The redox peaks increase as the number of cycle21 Page 8 of 25 Acc ep te d M an us cr ip t 8 increases, indicating additional electroactive ploy-beryllon II film deposition for each cycle. These 1 facts indicated that ploy-beryllon II membrane was deposited on the surface of CPE. After 2 electropolymerization, the modified electrode was carefully rinsed with doubly distilled water.3 Fig.14 3.2. Electrochemical characterization of different electrodes5 Fig. 2A shows the cyclic voltammetric responses of the bare CPE (a) and PBL-II/CPE (b) in 6 the 5.0 mM K3[Fe(CN)6] solution with 0.1 M KCl as the supporting electrolyte. The peak current 7 increased dramatically and the decrease of the peak-to-peak separation between the cathodic and 8 the anodic waves are clearly visible when the poly-beryllon II film is modified on the surface of 9 bare CPE. The results of the CVs demonstrate that the beryllon II film is conductive and does not 10 block electron transfer, which indicated that beryllon II modified CPE could greatly increase the 11 electron transfer rate of [Fe(CN)6] 3-/4-. For further characterization of the modified electrode and 12 clarify the differences among the electrochemical performance of bare CPE and PBL-II/CPE, 13 electrochemical impedance spectroscopy (EIS) was carried out. Fig. 2B shows impedance plots 14 for (a) bare CPE, (b) PBL-II/CPE in 5.0 mM K3Fe(CN)6/K4Fe(CN)6 (1:1) with 0.1 M KCl. The 15 semicircular elements correspond to the charge transfer resistances (Ret) at the electrode surface, a 16 large diameter was observed for the bar CPE in 79 k\u03a9 as curve shown. However, the diameter of 17 the semicircle diminished when PBL-II/CPE were employed. Curve b showed an arc, the diameter 18 of which displayed as Ret = 0.09 k\u03a9, which was really small than that for bare CPE. The charge 19 transfer resistance (Ret) values obtained from this observation implied that the charge transfer 20 resistance of the electrode surface decreased and the charge transfer rate increased using 21 PBL-II/CPE",
            "20 [25] Z.M. Liu, Z.L. Wang, Y.Y. Cao, Y.F. Jing, Y.L. Liu, Sens. Actuators. B. Chem. 157 (2011) 540.21 [26] K.J. Huang, L. Wang, Y.J. Liu, T. Gan, Y.M. Liu, L.L. Wang, Y. Fan, Electrochim. Acta 107 (2013) 379.22 23 24 25 26 27 28 29 30 31 32 33 34 Page 16 of 25 Acc ep te d M an us cr ip t 16 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 Figure captions:16 Fig. 1 displays the continuous CVs for the electrochemical polymerization of beryllon II over the 17 range of -2.0 to 2.0 V at 70 mV s\u22121 for 3 cycles.18 Fig. 2 A. CVs on bare CPE (a) and PBL-II/CPE (b) at the potential between \u22120.4V and 0.8V at a 19 scan rate of 100 mV s\u22121. B. Nyquist diagram of EIS at bare CPE (a) and PBL-II/CPE (b). EIS 20 condition: frequencies swept from 50 mHz to 100 kHz; signal amplitude: 0.005 V; solution: 5.0 21 mM Fe(CN)6 3\u2212/4\u2212 in 0.1 M KCl.22 Fig. 3 CVs of 50 \u03bcM PC at: bare CPE (a) and PBL-II/CPE (b) in 0.1 M PBS (pH 7.0) with the 23 scan rate of 100 mV s\u22121.24 Fig.4 A. CVs obtained at the PBL-II/CPE in 0.1 M PBS in pH values (a\u2192h:2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001853_s10846-015-0233-z-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001853_s10846-015-0233-z-Figure2-1.png",
        "caption": "Fig. 2 A multi-legged robot with six legs and 3-DOF per leg",
        "texts": [
            " The whole process is presented in a general algorithmic style for direct implementation in on-line and real-time controllers for multi-legged robots. The rest of the paper is organized as follows. Section 2 presents the modeling of a multi-legged robot. Section 3 presents the strategy of the reactive stability control method including an algorithm for a foot force based reactive stability control strategy. Section 4 presents the experimental demonstration of the presented methodology. The paper is concluded in Section 5. A general multi-legged robot is defined as a robot with n legs where each leg has m degrees of freedom (DOF). Figure 2 shows an example of a multi-legged robot that has six legs with 3-DOF per leg. Two coordinate frames are defined for the robot: B, the local body frame located on top of the robot body and G, the inertial frame located on the ground. The external force and moment perturbations are given by F and M , respectively. The presented reactive stability control method in this paper works for any multilegged robot as long as there is a solution to the inverse kinematics problem of the robot given by: IK : qij = f (O, R, ui, sij ) (1) where i = 1, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000195_tmag.2013.2245878-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000195_tmag.2013.2245878-Figure10-1.png",
        "caption": "Fig. 10. M1: Magnetic field distribution.",
        "texts": [
            " In Fig. 9, it can be known that the geometry of Section A is symmetric about the x-axis and the x coordinate of the acting point of the IUMP in Section A can be calculated by (31) Here, it is assumed that the acting point of local UMP of Section A is at (a,0). Because of the symmetrical structure of the stator, the action points of Sections B, C, and D are (0,a), (-a,0), and (0,-a), respectively. After the UMP in each segment is obtained by ANSOFT, the total UMP in one section can be calculated. Fig. 10 shows the magnetic field distribution of PMSM obtained with Maxwell 3-D static magnetic solver. From the analytical analysis in Section II, the order of fundamental harmonic of the UMP in space domain is 10, which is the same as the numerical results obtained with FEM shown in Figs. 11\u201313. Fig. 11 shows the IUMP component in the x direction, which is obtained with 3-D FEM. It can be found in the figure that the frequency of the x component is also ten times that of rotor rotating frequency, which is still the same as the one estimated with the analytical model in Section II"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000369_iros.2013.6696872-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000369_iros.2013.6696872-Figure2-1.png",
        "caption": "Fig. 2. The biped configuration and the impulsive actuations at the end of swing. The leg transitioning from swing to stance is thick red. a) At the end of single stance, when \u03c6=2\u03b1, an impulsive force is applied along the stance leg by the leg prismatic actuator (not shown in the figure), and an impulsive retracting (contracting the hip angle) torque is applied by the hip actuator. P and R are the impulse intensity (time integral) of these force and torque. b) The swing foot (F2) touches the ground immediately after push-off impulse P and retraction impulse R are completed.",
        "texts": [
            " For the sake of simplicity, we approximate the burst push-off force and the burst retraction torque as theoretical impulsive (infinitesimal duration with infinite magnitude) force and torque that cause discontinuous velocity jumps in an exactly fixed biped configuration. Although this approximation is unrealistic, the insights achieved from the analysis of the resulting simplified models are an important step in improving our understanding of effective legged locomotion. For example, the impulsive models have been previously used to explore the energetic benefits of pre-emptive pushoff in bipedal walking [8], [13] and the foot sequencing in animal gaits [8], and the energetic consequence of step-tostep transitions in human walking [14]. Fig. 2a shows the biped at the end of swing phase (single stance) when the impulsive push-off force and the impulsive retraction torque are applied on the biped during an infinitesimal period just before heel-strike. The impulsive push-off force and the impulsive retraction torque are quantified by their net impulse denoted by P and R. If we denote the time instant just before both impulsive actions by t\u2212pr and the time instant just before heel-strike (just after both impulsive actions are completed) by t\u2212h , the push-off impulse P and the retraction impulse R are given by P = \u222b t\u2212 h t\u2212pr F (t) dt (1) R = \u2212 \u222b t\u2212 h t\u2212pr \u03c4(t) dt. (2) We assume that the sign of the push-off force and the retraction torque does not change during their application. Therefore, based on our convention in Fig. 1 for the positive directions, the push-off force F is always positive, and the retraction torque \u03c4 is always negative. Hence, P > 0 and R> 0. Although the impulsive push-off and retraction actions take place at the same biped configuration (Fig. 2a), we treat P and R as isolated in time (one after the other) or as having some (specified) overlap with each other (to be discussed later in detail). We will show in section IV that the relative timing of these two impulses can change the energetics of the gait. Immediately after both impulsive actions are completed the swing foot hits the ground (Fig. 2b). Our analysis in this paper does not depend on whether heel-strike is assumed to be impulsive or non-impulsive. Our focus in this section is on the final moment of the swing phase, when the impulsive push-off force and the impulsive retraction torque are applied on the biped just before heel-strike. We use t\u2212p and t+p to denote the time instants just before and just after the impulsive push-off P , and t\u2212r and t+r to denote the time instants just before and just after the impulsive retraction R"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001616_physreve.98.052141-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001616_physreve.98.052141-Figure1-1.png",
        "caption": "FIG. 1. (a) Lattice used for models 1, 2, and 3. Particles can move in any of the four directions indicated by the arrows. (b) Lattice used for model 4. Here, six directions are possible. (c) RTP model, where particles move in straight lines between tumble events. (d) Particles in the ABP model change direction continuously. In all cases, the wall is designed by a thick vertical line.",
        "texts": [
            " The position update of the particles is asynchronous to avoid two particles attempting to jump to the same site simultaneously: at each time step, the N particles are sorted randomly and, sequentially, each particle attempts to jump. After the position updates, with some small probability \u03b1, independently for each particle, tumbles take place and the sk changes to a new random direction [37]. The four variations of the model differ in the lattice geometry, in the form the maximum occupancy is enforced, and the way the new directions are chosen after a tumble. This case corresponds to the original PEP model. Here, the lattice is square, composed of Lx \u00d7 Ly sites and, consistently, sk = {\u00b1x\u0302,\u00b1y\u0302} [see Fig. 1(a)]. If the occupation of the destination site is smaller than the maximum occupancy per site nmax, the jump takes place; otherwise, the particle remains at the original position. That is, the maximum occupancy is strictly enforced. For tumbles, the directors sk are redrawn at random from the four possibilities, independently of the original value. Here, the lattice and the tumble protocol are the same as in model 1. To describe swimmers or agents that can deform or be compressed, the jumps of the particles to the new site are now probabilistic, depending on occupation fraction of the destination site nrk+sk /nmax, where rk is the position of the kth particle",
            " Values and ranges used for the parameters in the simulations of the ten models under study. Model \u03bd\u0302T nmax V\u03020 \u03c6 Lx Ly Wall 4 [0.001,0.03] 2, 3 \u2217 0.040 3000 500 \u2217 5 [0.002,0.20] \u2217 0.0002, 2 0.004 6000 1000 A 6 [0.002,0.20] \u2217 2 0.004 6000 1000 A 7 [0.002,0.20] \u2217 0.2 0.010 6000 1000 A 8 [0.002,0.20] \u2217 0.2 0.010 6000 1000 A 9 [0.002,0.20] \u2217 0.0002, 2 0.010 6000 1000 B 10 [0.002,0.20] \u2217 2 0.010 6000 1000 C 052141-2 Finally, the lattice is changed to a triangular one with a major axis parallel to the wall [see Fig. 1(b)]. The six values of the state variable are sk = {\u00b1y\u0302,\u00b1 \u221a 3 2 x\u0302 \u00b1 1 2 y\u0302}. Maximum occupancy is enforced as in model 1 and the tumbles have no memory, but now the new direction is chosen with equal probability between the six directions. For the off-lattice cases we consider the run and tumble particle (RTP) and the active Brownian particle (ABP) models of active particles [3,4]. A total of N particles move in a continuous 2D surface of area Lx \u00d7 Ly and time evolves continuously. Each particle is characterized by its position rk and director sk = (cos \u03b8k, sin \u03b8k ), with \u03b8k \u2208 R",
            " 4 or 5), contrary to the Gaussian potential, where particles can occupy the same position (see first row of Fig. 3). The Gaussian potential is appropriate for geometries where particles can move partially in the third dimension and, hence, appear to overlap in the plane. Particle-particle interactions do no modify explicitly their propulsion speed neither they align the directors as in Vicseklike models. Indeed, directors evolve independently for each particle. For the RTP model, with a rate \u03b1, particles can experience a tumble [see Fig. 1(c)]. In an instantaneous event, 052141-3 the direction changes to a new one, s\u2032 k , which is chosen from a kernel W (sk, s\u2032 k ) that could depend on the original director. For isotropic media W (sk, s\u2032 k ) = w(sk \u00b7 s\u2032 k ) and we take W = 1/2\u03c0 , indicating that the new director is chosen completely at random. In the case of the ABP model [see Fig. 1(d)], the director sk changes continuously as an effect of rotational noise, a process that is described with a Langevin equation, dsk dt = \u221a 2Dr sk \u00d7 \u03b7k (t ), (5) where Dr is an effective rotational diffusion coefficient and \u03b7k = \u03b7k z\u0302 (z\u0302 is perpendicular to the plane where the particles move) is a white noise of zero mean and correlation \u3008\u03b7k (t )\u03b7k\u2032 (t \u2032)\u3009 = \u03b4kk\u2032\u03b4(t \u2212 t \u2032), where k and k\u2032 label the particles. Equation (5) is equivalent to \u03b8\u0307k = \u221a 2Dr\u03b7k (t ) in two dimensions. The wall is modeled in three different ways"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002210_1.4029828-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002210_1.4029828-Figure3-1.png",
        "caption": "Fig. 3 SE model of a rotating ring",
        "texts": [
            " The numerical validation of the rotating ring spectral element derived in this paper is done for a uniform homogeneous rotating ring by comparison of the forced response and operational modes with the analytical solution. The properties of the ring are: mass density q\u00bc 7200 kg/m3, Young\u2019s modulus E\u00bc 220 GPa, radius R\u00bc 0.25 m, rectangular cross section with width b\u00bc 0.15 m, height h\u00bc 0.002 m, area A\u00bc bh, and area moment of inertia I\u00bc bh3/12. Arbitrarily, the spectral element (SE) model consists of four spectral elements (Fig. 3) (for a uniform ring one spectral element would suffice). The analytical solution, which includes the effects of the internal pressure and/or an elastic foundation (see Appendix) is used to compute natural frequencies, mode shapes, and the forced response. 3.1 Rotating Effects: Coriolis Acceleration and Centrifugal Hoop Stress. In this section, the ring rotation effects are investigated. For the sake of clarity, the effects of internal pressure, elastic foundation, and structural damping are neglected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002808_s11071-016-3072-y-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002808_s11071-016-3072-y-Figure5-1.png",
        "caption": "Fig. 5 Geometric relation between the manipulator and the gun",
        "texts": [
            "01), we use the IL control. Otherwise, we use traditional PD control, the parameters of which is derived from the IL controller of the previous step. 3.4 Desired trajectory The purpose of the tracking control of the ammunition transfermanipulator is to realize automatic loading while the gun is pitching, so it is important to configure the relation between the trajectory of the manipulator and the angle position of the gun. The geometric relation between the manipulator and the gun in loading position is presented in Fig. 5, where B4 denotes the gun, O4 is the trunnion axis, \u03b8 is the angle of the gun, other notations are similar with which in Fig. 5. It is easily to obtain following equations yr2 = yO4 \u2212 xO4tan\u03b8 \u03b83 = \u03b8 Here, we give the guns pitching trajectory as \u03b8 = sin(\u03c9t) Then, we could obtain the equations of desired trajectory for the manipulator as follow y\u0307r2 = \u2212xO4\u03c9cos(\u03c9t) \u2212 xO4\u03c9cos(\u03c9t)tan2 (sin(\u03c9t))\u03b8\u03073 = \u03c9cos(\u03c9t) In this section, in order to evaluate the proposed control system, simulations are demonstrated byMATLAB software, with parameter setting as shown in Table 1. The oscillation of the base was obtained by virtual prototype simulations of a tracked vehicle using ADMAS/ATV software, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000935_cbo9781139013451.017-Figure13.4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000935_cbo9781139013451.017-Figure13.4-1.png",
        "caption": "Figure 13.4. A third-grader\u2019s design plans and corresponding three-pitch musical instrument from the curriculum\u2019s overarching engineering design challenge.",
        "texts": [
            " Students explore the structural design of these instruments, observe how they look and sound when played, and consider how visible characteristics are related to sound characteristics. For example, how does the size of an object influence the pitch of the sound it makes? Throughout the unit, students are encouraged to consider how these relationships can inform their design of a new musical instrument. In the unit\u2019s two concluding sessions, students design, construct, and demonstrate musical instruments of their own invention (see Figure 13.4). They also demonstrate what they have learned about the unit\u2019s big science question: How are sounds made? The teacher\u2019s guide for the Design a Musical Instrument unit is intended both to guide the teacher\u2019s facilitation of the learning and to support growth in the teacher\u2019s science and pedagogical content knowledge. It suggests that teachers and students follow a similar sequence of events in each individual session. This begins with introducing the scientific question to be investigated. Next, at https:/www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001537_0954406218791636-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001537_0954406218791636-Figure11-1.png",
        "caption": "Figure 11. Representation of the turbine-generator shaft line layout after alignment and displacement-correction of the bearing shell position. Rotor centreline approximates the casing centreline but the journal retains its position inside the bearing\u2019s clearance.",
        "texts": [
            " With the proper corrections in bearing pad centre, the journal centring to the casing can be optimized. The loading at each bearing does not change considerably because the entire shaft line is lifted in relatively similar amount at each bearing location. This is the reason that the relative bearing eccentricity of operation does not change as well. The journal orbit retains its characteristics of extent (amplitude). The system is supposed now to operate with the rotor\u2019s centreline well approximating the casing\u2019s centreline, see Figure 11. The bearing performance (e.g. regarding power loss) does not change as well, see Table 5. The alignment configuration is negligibly changed and therefore the static load at each bearing remains approximately the same as in conventional bearing operation, see Table 5. In this section, the adjustable bearing is demonstrated regarding its ability to optimize its fluid film performance concerning the resulting friction coefficient. The analysis concerns steady-state operation at the rated speed of the turbine-generator shown in Figure 6, which is still defined at 3000 r/min as in previous examples"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000754_j.petlm.2015.10.012-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000754_j.petlm.2015.10.012-Figure1-1.png",
        "caption": "Fig. 1. Finite element model.",
        "texts": [
            "2015.10.012 modeling. In addition, because of the symmetry of the bearing structure, 1/2 of the model was constructed and analyzed, and the model which is consisting of the three teeth respectively taking from the three gear rings of cone bit was created at the same time. Those teeth can meet demands of the force analysis. Model was meshed with the SOLID187 element, and the element size was 3 mm. The finite element model of hollow cylindrical roller bearing of cone bit after meshing is shown as Fig.1. Besides, the hollow size of the roller is 40%, and there are three contact pairs in the hollow cylindrical roller bearing, thrust bearing and the path to the sliding bearing respectively. The friction coefficient is 0.05 on the contact pairs, and the initial contact interference is 0.001 mm. The material of cone is 20Ni4Mo, and the material of claw journal is 20CrNiMo, and the material of cylindrical roller is 55SiMoV. Elastic modulus is 218 GPa, 219 GPa and 210 GPa respectively, and Poisson's ratio is 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001707_042148-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001707_042148-Figure6-1.png",
        "caption": "Figure 6. Structures of the flexspline.",
        "texts": [],
        "surrounding_texts": [
            "The HYPERMESH was used for the pretreatment of solid model, including mesh, add material attributes and determine boundary conditions according to actual work conditions. Then, the K file was imported into the LS_DYNA to do the finite element analysis. The specific parameters were shown in Table 3. The FEA result of stress distribution of the failed flexspline meshing parts was shown in Fig. 7. The stress was indicated in Mpa units. According to the FEA, the weakest regions were tooth root which near the working surfaces. The maximum stress of the failed flexspline calculated by FEA was 162.683 Mpa. IMMAEE 2018 IOP Conf. Series: Materials Science and Engineering452 (2018) 042148 IOP Publishing doi:10.1088/1757-899X/452/4/042148 By comparing the microstructure of the material, it was found that there are cracks at the tooth root surface, where the stress concentration and the alternating stress are the root causes of the failure of the flexspline. In order to better analyze the stress field of the failed flexspline, the influence of different fillet radii at bending part on the maximum stress of the flexspline under no load was studied by taking the failed flexspline as an analysis model. The fillet radius is 0~2 mm, and the other parameters are unchanged. The FEA result of stress distribution is shown in Fig. 8. Fig. 8a shows that, when the flexspline is running, the force at bending part is the largest, and the radii of the fillet here directly affect the maximum stress of the flexspline. Fig. 8b shows the maximum stress of the flexspline corresponding to different fillet radii: 1) when the radius is 0 mm, the maximum stress of the flexspline is the largest, reaching 714.7 Mpa; 2) when the radius is 0.2 mm, the maximum stress of the flexspline is 137.9 Mpa, and as the radius increases, the maximum stress of the flexspline decreases; 3) when the radius is about 0.8 mm, the maximum stress of the flexspline reaches the minimum, 94.2Mpa, with the radius increasing, the maximum stress of the flexspline increases; 4) when the radius is 2 mm, the maximum stress of the flexspline is 149.3 Mpa. Taken together, when the fillet radius is about 0.8 mm, the stress is minimum at bending part. By changing the fillet radius, the stress distribution of the flexspline, in a manner, can be improved."
        ]
    },
    {
        "image_filename": "designv11_13_0000513_smc.2014.6974513-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000513_smc.2014.6974513-Figure2-1.png",
        "caption": "Figure 2 The definition of the arm plane",
        "texts": [
            " The D-H frames and This work was partly supported by the National Natural Science Foundation of China under Grant 61175098, and 51205078 and Aerospace Science and Technology Innovation Fund of China under Grant CASC201102. 978-1-4799-3840-7/14/$31.00 \u00a92014 IEEE 3744 parameters are respectively shown in Figure 1 and listed in A. Arm Plane and Arm Angle According to the characteristics of a S-R-S manipulator, three joint axis of the shoulder (joint 1~3) and the wrist (joint 5~7) form an equivalent spherical joint. The intersection points are denoted by the point S and the point W respectively. In addition, the origin of the 3th frame which is located on the elbow (joint 4) is denoted by the point E. As shown in Figure 2, the plane SEW is defined as the arm plane [8]. Given a reference plane, the arm angle \u03c8 , i.e. the angle between the reference plane and the arm plane, is defined to represent the self-motion [8]. Without loss of generality, the line SW is used as a line of the reference plane. Then only one more point Q is needed to determine the reference plane, which is then represented as SQW. The relationship of the reference plane, the arm plane and the arm angle is shown in Figure 3. For the convenience of discussion, we define some vectors as follows: w : A vector from the point S to the point W; e : A vector from the point S to the point E; V : The vector from the point S to the point Q in the reference plane; k p, : The vectors which are perpendicular to w in the B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000053_1464419314566086-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000053_1464419314566086-Figure2-1.png",
        "caption": "Figure 2. Roller raceway contact in a SRB.",
        "texts": [
            " It should be stated that this assumption is idealized and not true in practice. 4. The effect of centrifugal force on roller load is ignored. 5. The effect of lubrication is ignored. However, this is not the case in practice and some analysis of lubricated contact dynamics has been carried out.30,31 The present study essentially follows the Hertzian dry load approach. Contact mechanics analysis of roller raceway In elastic contact range, the contact of roller raceway in the SRB can be considered Hertzian elliptical contact,12 as shown in Figure 2. Fr is the load acting on the inner ring; Q1 j and Q2 j are the roller loads at the azimuth angle j of the two rows in the SRB, respectively. It is assumed that the rollers do not misalign much, and the roller-to-raceway contact in the SRBs yields elliptical contact footprint shape. By using the curvefitting method, Houpert5 derived a set of formulae for calculating Hertzian elliptical contact parameters for SRBs, which are expressed as a Rx \u00bc 1:5528k0:3737W1=3 PC \u00f01\u00de b Rx \u00bc 1:1063k 0:1866W1=3 PC \u00f02\u00de Rx \u00bc 1:7138k 0:2743W2=3 PC \u00f03\u00de where a and b are the semi-length in the transverse direction and semi-width in the direction of rolling in the elliptical contact, respectively; k and WPC are dimensionless parameters that can be expressed as k \u00bc Ry Rx \u00f04\u00de WPC \u00bc Q E0R2 x \u00f05\u00de Substituting equations (4) and (5) into equation (3) yields \u00bc 1:7138k 0:2743Rx 1 E0R2 x 2=3 Q2=3 \u00f06\u00de When one roller contacts with the inner and outer raceways simultaneously, the following equations can be obtained32 i \u00bc GiQ 2=3 \u00f07\u00de e \u00bc GeQ 2=3 \u00f08\u00de In equations (7) and (8), Gi and Ge are flexibility coefficients in the contact regions, and can be calculated by the following equations Gi \u00bc 1:7138k 0:2743i Rix 1 E0R2 ix 2=3 \u00f09\u00de Ge \u00bc 1:7138k 0:2743e Rex 1 E0R2 ex 2=3 \u00f010\u00de In equations (9) and (10), E0 is the equivalent elastic modulus of the two contact bodies, and its value can be obtained by 1 E0 \u00bc 1 21 E1 \u00fe 1 22 E2 \u00f011\u00de For the steel rolling element bearings, E1 \u00bc E2 \u00bc 2:07 105MPa, 1 \u00bc 2 \u00bc 0:3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001777_978-1-4419-8420-3-Figure2.10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001777_978-1-4419-8420-3-Figure2.10-1.png",
        "caption": "Fig. 2.10 Guarding and inspection process on a more complicated pipeline. a Upper simulated model. Lower guarding points and guarded regions rendered in different colors, respectively. b Damaged pipe with big and concave holes. c Large holes are detected/extracted from more than one guarding points; the whole big boundary is composed of several extracted subboundaries and identified separately from different guarding points, as shown in green",
        "texts": [
            " Any given region of the pipeline is covered by at least one guard and, therefore, is colorized. The height maps can then be generated as templates, measuring the \u201ccorrect\u201d distance from each guarding site to the pipe wall toward specific directions. This simulates range images obtained by a laser scanner. Now, we simulate the appearance of defected regions on the pipeline by generating some missing regions as shown in (b). When the robot checks height maps on guarding points, these holes can be immediately detected and illustrated in (c). Another example is shown in Fig. 2.10; this pipe is guarded by 12 points (a). In addition, the damaged region of the pipe is big and with complex topology (b). In this case, the robot should check from more than one guarding points in order to detect the entire shape of such a big hole. The entire defect geometry is extracted by composing boundaries detected from different guarding sites. The merged boundary loop is illustrated in (c). 2.5 Result of Simulated Experiments 33 Small deformation such as bending and erosion can also be detected in our system as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000536_worv.2013.6521918-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000536_worv.2013.6521918-Figure2-1.png",
        "caption": "Figure 2. Mechanical compliance.",
        "texts": [
            " There are an LRF fixed above food on a plate, a manipulator, a spoon attached at the tip of manipulator and food on a plate. In order to reduce the noise with measurement, a plate and a table that lack luster are used. A plate is not fixed on a table, which implies that the plate might move slightly by each scooping motion. The LRF and manipulator are connected to a PC. The LRF sends 3D point data to PC. The command sent from PC to manipulator provides trajectories of the tip of the manipulator. Two springs are attached between the spoon and the hand tip, as depicted in Fig. 2. This mechanical compliance can soften the force between spoon and plate. The work flow for scooping is shown in Fig. 3. Before scooping work, the calibration of coordinate transformation between LRF and manipulator is conducted following the method using quaternions [11]. The first step in the flow is 3D measurement of a food using LRF. This measurement by LRF is executed before each scooping motion. The obtained set of 3D points includes those corresponding to the plate. By subtracting points of the plate measured in advance, points corresponding to the food are extracted as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000310_j.humov.2012.07.003-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000310_j.humov.2012.07.003-Figure1-1.png",
        "caption": "Fig. 1. Illustration of the calculated parameters. The reflective markers are illustrated by open circles.",
        "texts": [
            "0 , which was marked on the floor. A valid kick had to hit a target area of 1 by 1 m, starting 0.15 m above the floor, which was attached to a net 4 m away from the ball. The order of testing was randomized between subjects. Three spherical reflective markers were attached in a triangle to each of the two balls. Four other markers were attached to the hip (troachanter major), knee (condylus lateralis femoris), ankle (malleolus lateralis), and toe (caput metatarsalis V) of the kicking leg of each participant (see Fig. 1). Threedimensional motion of the markers was recorded with an eight-camera opto-electronic system operating at a frame rate of 500 Hz (ProReflex MCU1000, Qualisys Medical AB, Gothenburg, Sweden). The geometrical center of the ball was calculated from the three markers placed on the surface of the ball (Andersen, Kristensen, & S\u00f8rensen, 2008).The beginning of the impact was determined as the first frame, where the ball showed positive acceleration. Vball was defined as the average resultant velocity over 10 frames, starting 16 ms (8 frames) after the beginning of the impact, since the duration of the impact period is between 10 and 13 ms (Kellis & Kattis, 2007; Lees et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000739_s0263574713000829-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000739_s0263574713000829-Figure2-1.png",
        "caption": "Fig. 2. Supporting convex hull.",
        "texts": [
            " Then, joint torques are calculated by \u23a1 \u23a306\u00d71 \u03c4rl \u03c4ll \u23a4 \u23a6 = Mq\u0308 + C \u2212 WJT rf Kf W f\u0302ext \u2212 WJT lf (I \u2212 Kf )W f\u0302ext. (15) The first six torques are zeros because they are virtual joints and do not have kinematic constraints on the floating base. Kf is the force distribution matrix. Desired accelerations are obtained by (14). Their corresponding external forces are obtained by (12) and (13). The desired external forces are not always equal to the GRFs because the bipedal robot is an unactuated system and the supporting convex hull (SCH) is limited (see Fig. 2). If the desired ZMP exceeds the SCH, the ankles would have an unexpected rotation. Friction is related to the vertical force and friction coefficient; so the desired friction, divided by the desired vertical force, must be smaller than the friction coefficient. If it is not, the feet of the robot slip. The ground can only provide the unilateral force vertically, so the vertical force must be positive. The ground can only push the feet. All in all, actual external forces satisfy these constraints",
            " First, the relationship between the accelerations of the floating base and their corresponding external forces is deduced. Rewrite M(q) as follows: M(q) = \u23a1 \u23a3M11 M12 M13 M21 M22 M23 M31 M32 M33 \u23a4 \u23a6. Each element of this matrix is an R6\u00d76 matrix. The following equation is obtained by (7): M11q\u0308r + M12q\u0308rl + M13q\u0308ll + C1 = WJT M1 MXT W Mf\u0302m, (16) http://journals.cambridge.org Downloaded: 28 Jun 2014 IP address: 155.198.30.43 where MXW is the coordinate transformation matrix from \u2211 W to frame \u2211 M (see Fig. 2). \u2211 M is selected to locate where GRFs provide the maximum or minimum accelerations. Mf\u0302m is the desired external force expressed in this frame. Equation (16) is simplified as follows by (9) and (10): Hq\u0308r + D = WMf\u0302m, (17) where W = WJT M1 MXT W , (18) H = M11 \u2212 M12 WJ\u22121 rf2 WJrf1 \u2212 M13 WJ\u22121 lf3 WJlf1, (19) D = C1 \u2212 M12 WJ\u22121 rf2 W J\u0307rf q\u0307 \u2212 M13 WJ\u22121 lf3 W J\u0307lf q\u0307. (20) The following equation is obtained by (8), (11) and (17): Uac + V = Mf\u0302m, (21) where U = W\u22121HWJ\u22121 R RX\u22121 W , (22) V = W\u22121 ( D \u2212 HWJ\u22121 R ( RX\u22121 W [03\u00d71; R\u03c9r \u00d7 Rvr ] + W J\u0307Rq\u0307r )) ",
            " Choosing \u2265 or \u2264 depends on where the coordinate frame is. For instance, if the frame is at the front side of the SCH, \u2265 is selected; if the frame is at the back side of the SCH, \u2264 is selected. nMx \u2265 or \u2264 0 guarantees the desired ZMP within the SCH in the y direction. If the frame is at the left side of the SCH, \u2264 is selected; if the frame is at the right side of the SCH, \u2265 is selected. http://journals.cambridge.org Downloaded: 28 Jun 2014 IP address: 155.198.30.43 There are two constrained Mf\u0302m which are expressed in two \u2211 M (see Fig. 2), respectively. According to the force constraints mentioned above and (21), we can obtain five inequalities for each Mf\u0302m. These 10 inequalities are related to U , V and Mf\u0302m. We rewrite them into matrix form as A \u2217 ac act \u2264 b, (24) where A \u2208 R10\u00d73 and b \u2208 R10\u00d71 are functions of U , V and Mf\u0302m, respectively, ac act is the actual acceleration which should satisfy its constraints. When we get desired acceleration ac des by (14), QP is used to find extremal ones. The cost function is f (ac act) = ( ac act \u2212 ac des )T \u2217 Wt \u2217 ( ac act \u2212 ac des ) , (25) where Wt is a weight matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003460_s12206-018-0441-0-Figure8-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003460_s12206-018-0441-0-Figure8-1.png",
        "caption": "Fig. 8. Limitation of the Paden\u2013Kahan definition.",
        "texts": [
            " 0 st st e e e e e e e g g G q \u03b8 \u03b8               = = $ $ $$ $ $ (24) \u03b84 and \u03b85 are solved using subproblem 2, where p = qe, q = G2(\u03b8)qe and r = qw. The last rotation angle \u03b86 is simply obtained using subproblem 1. ( ) ( ) ( ) 6 6 3 3 5 51 1 2 2 4 4 3 1 [$ ]\u03b8 [$ ]\u03b8 [$ ]\u03b8[$ ]\u03b8 [$ ]\u03b8 [$ ]\u03b8 -1 st = \u03b8 g (0) \u03b8 . r r e st e p e e e e e g p G p= - (25) In this example, pr is a random point that does not exist in S6 and G3(\u03b8) is transform matrix between end-effector pose and last rotation angle. Subproblems 2 and 3 are not intuitive for a robot beginner, and the two skew axis vectors, such as those shown in Fig. 8, are not solved because two screw axis are not intersection. In this section, a new subproblem is introduced to solve the inverse kinematics when the two axis vectors are skewed in space with a zero pitch and a unit magnitude. New subproblems include the subproblems 2 and 3, as shown in Fig. 9. A new subproblem provides a solution of inverse kinematics for all cases in space and is geometrically explained using skew coordinates. 4.1 Introduction of skew coordinates In this section, skew coordinates are introduced to define a new subproblem"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000767_tro.2015.2482078-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000767_tro.2015.2482078-Figure1-1.png",
        "caption": "Fig. 1. Model of the environment showing the robot of radius r at location x. A sink element is located at the destination location xd . To guide the robot on a nonholonomic trajectory, a source element is placed at a distance \u0394d behind the robot at location xs .",
        "texts": [
            " (1) For example, the potential function for a point charge located at p = [px py pz ]T is defined as \u03c6(x) = \u2212 Q 4\u03c0 \u2016x \u2212 p\u2016 (2) where Q = 0 represents the strength of the charge. It can be shown that the potential function in (2) is an HPF since \u22072\u03c6(x) = 0. The point charge potential function can be used to model both source (repulsion) and sink (attraction) elements, where Q > 0 and Q < 0 represent source and sink point charges, respectively. The model of the environment for the proposed 3-D motion planner is shown in Fig. 1. The robot\u2019s physical dimensions are considered by assuming that the robot can be encapsulated within a sphere of radius r. This radius can be considered when avoiding obstacles by incorporating it into the radius of the obstacles. The pose of the robot can be defined by the position of its center x = [x y z]T expressed in Cartesian coordinates and its orientation q = [qx qy qz \u03b8]T represented in quaternion form. The heading direction of the robot d can be defined by the unit vector d = q q0 q\u22121 (3) where q0 = [1 0 0 0]T represents the inertial frame of the reference quaternion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001373_j.triboint.2018.01.034-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001373_j.triboint.2018.01.034-Figure3-1.png",
        "caption": "Fig. 3. Simulation model of oil film between the disks of HVD.",
        "texts": [
            " The energy equation for the flow can be written as follows \u03c1cp vr \u2202T\u00fev\u03b8 \u2202T \u00bck \u22022T 2 \u00fe1 \u2202T\u00fe 1 2 \u22022T 2 \u00fe \u22022T 2 \u00fe\u03bc \u2202v\u03b8 2 \u00fe \u2202vr 2 \u2202r r \u2202\u03b8 \u2202r r \u2202r r \u2202\u03b8 \u2202z \" \u2202z \u2202z # (33) where T is the fluid temperature, cp is the specific heat at constant pressure, k is the thermal conductivity of the oil film. Isothermal boundary conditions for Eq. (33) are as follows [21]: T\u00f0r; \u03b8; 0\u00de \u00bc T\u00f0r; \u03b8; h\u00de \u00bc T\u00f0a; \u03b8; z\u00de \u00bc Tin \u2202T \u2202r jr \u00bc b \u00bc 0 (34) where Tin is the oil temperature at the inner radius. Since the simplified momentum equations couple with energy equation, it is almost impossible to solve the equations above analytically. Therefore the flow is solved numerically by means of computational fluid dynamics (CFD) code FLUENT based on Finite Volume Method in this case. Fig. 3 shows the numerical simulation model about the oil film between the disks. The gray area indicates the fluid computational (b) Interface with radial deformation.Fig. 16. Effect of radial deformation at the inlet and outlet on viscous torque Th when t\u00bc 0.5s. domain. The two identical annular discs with inner radius a and outer radius b are separated by a distance h along the direction of the rotation axis z. The left disk and the right disk rotate with angular velocity \u03c92 and \u03c91, respectively. Due to the non-slip condition at the surface of the disks, the rotation of the disks sets the fluid in the gap in motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003073_060201-Figure4-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003073_060201-Figure4-1.png",
        "caption": "Fig. 4. (color online) Coupling design of nickel-based superalloy electrode conical surface and induction coil.",
        "texts": [
            " The macroscopic magnetic field can be described by the Maxwell equations as follows: \u2207\u00d7H = J+ \u2202D \u2202 t , (6) \u2207\u00d7E =\u2212\u2202B \u2202 t , (7) \u2207 \u00b7B = 0, (8) \u2207 \u00b7D = \u03c1, (9) where H refers to the magnetic field intensity (in units A/m), J the current density (in units A/m2), D the electric flux density (in units C/m2), E the electric intensity (in units V/m), B the magnetic flux density (in unit T), \u03c1 the electric density (in units C/m) and t the time (in unit s). These expressions describe the relationship among electromagnetic field quantities (H, D, E, B) in some domain, as a function of field source electric charge and electric current. The coupled design of the superalloy electrode tip shape and conical induction coil is critical for the induction melting process. Figure 4 shows the details of the superalloy electrode tip shape and induction coil in a consecutive induction melting experiment described in Subsection 3.1. The coupled design is further verified using the finite element software COMSOL Multiphysicsr. As shown in Fig. 5, the electromagnetic field energy concentrates on the conically shaped electrode, which is immersed in a tapered coil. The current flows in a thin skin and concentrates on the inner surfaces of the coil due to magnetic flux confinement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002722_012101-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002722_012101-Figure5-1.png",
        "caption": "Figure 5. Results of the dynamic analysis.",
        "texts": [
            " The dynamic analysis of the platform movement was carried out in order to determine the maximal forces that are applied to the components of the platform (forces actions in particular joints). In order to analyze the obtained results they were exported to the Microsoft Excel program, where they were processed. First was selected the leg of the platform, in relation to which the loads, during the simulation, were the highest. And then was selected this step of the dynamic analysis during which was observed the maximal loads of the analyzed components of the platform. In the figure 5 are shown the simulation results of the analyzed movement of the platform. The results, obtained in this way, were then the basis for conducting the strength analysis with the FEM method. 4. Strength analysis in the Advanced Simulation module The last stage of the realized research cycle was carrying out the strength analysis with the FEM method. This analysis was conducted in the Advanced Simulation module using the NX Nastran solver. It was analyzed this leg of the parallel robot in which occurred the highest loads"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002411_ssd.2015.7348200-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002411_ssd.2015.7348200-Figure1-1.png",
        "caption": "Fig. 1. Experimental setup.",
        "texts": [
            " The proposed algorithm has been applied on plug/socket type IEC 60309 which is very hard for plugging, even for the human. The paper is organized as follows. In the next section, the experimental setup will be presented. After this, section III describes the proposed algorithm the of peg-in-hole task. In section IV the implementation of the proposed algorithm in a commercial controller and the experimental results will presented. The last section contains discussion and conclusion. The experimental setup can be seen in Fig. 1. It consists of the six-axes articulated robot KUKA KR6/2 which is 978-1-4799-1758-7/15/$31.00 \u00a92015 IEEE controlled by the industrial robot controller of type KRC2. The robot is equipped with the FT Delta SI-660-60 force sensor by SCHUNK. It is a six component force/torque sensor with the effective measurement range of \u00b1660N and \u00b160Nm for forces and torques, respectively. The sensor is mounted on the manipulator flange and it is protected by a pneumatic robot load limiter. The robot end-effector consists of a two-fingerparallel gripper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003272_ssd.2017.8166915-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003272_ssd.2017.8166915-Figure1-1.png",
        "caption": "Fig. 1. Architecture of the control.",
        "texts": [
            " INTRODUCTION Recently, energy consumption has increased enormously, due to massive industrialization, that\u2019s why the exploitation and development of renewable energies have grown strongly. Today, wind power has become a reliable Source for energy production due to its motivating nature, the evolution of semiconductor technology and the new methods of controlling the variable speed turbines [1]. Wind turbines based on DFIG whose stator is directly connected to the grid contrary to the rotor which is connected via power converters (fig.1) are nowadays the most widely used because it allows producing electricity at variable speed and thus to better exploiting the wind resources for different wind conditions [2]. But, the major difficulty in controlling this type of machine lies in the fact that the powers are strongly coupled. For this reason, several methods of controlling of DFIG have emerged, among them ,the field oriented control technique .The principle of this technique was developed by BLASCHKE in the early 1970s; it consists in orienting the field along one of the axes of the referential d, q in order to make this machine\u2019s behavior similar to that of a Separately Excited DC Machine, this control is based on PID controllers In this article, we begin with the modeling of the turbine [3][4], and next we present a model of the DFIG in the referential (d, q). After that, we present the principle of direct and indirect stator field oriented control. Finally with the use of MATLAB / SIMULINK, we will present and analyze the simulation results to validate our theoretical study. The general structure of the wind system and control FOC is given in ( fig. 1) [12]: II. MODELING OF THE CONVERSION CHAIN The wind power is given by: = . . (1) The turbine is a device to convert a percentage of the wind power, defined by Cp, so aerodynamic power of the turbine will be given by:= (\u03bb, ). . . . (2) 978-1-5386-3175-1/17/$31.00 \u00a92017 IEEE The speed ratio \u03bb [13], which is defined as the ratio between the linear speed of the turbine \u03a9t and the wind speed v:\u03bb = R. (3) The aerodynamic torque is defined from the speed of the turbine as follow: = (\u03bb, ). . . . . (4) The power coefficient Cp is approximated using a nonlinear equation described as follows [5]:C ((\u03bb, \u03b2) = C \u2212 C "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001294_tpel.2017.2782804-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001294_tpel.2017.2782804-Figure6-1.png",
        "caption": "Fig. 6. Vector sum in the air gap of six phase IPMSM: (a) current sum, and (b) flux sum.",
        "texts": [
            "org/publications_standards/publications/rights/index.html for more information. This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication. Citation information: DOI 10.1109/TPEL.2017.2782804, IEEE Transactions on Power Electronics 11 In the six-phase IPMSM, we have two current vectors in the air gap; iedq1 and iedq2. But the latter is \u03c0 6 advanced from the former. Therefore, the air gap current vector is iedq1 + ej \u03c0 6 iedq2 as shown in Fig. 6 (a). Similarly, it is necessary to consider the flux sum \u03bbg in the air gap (Fig. 6 (b)) as \u03bbg = (Ldq1iedq1 +\u03c8edq1) + ej \u03c0 6 (Ldq2iedq2). (34) Therefore, the torque of P -pole six phase IPMSM is equal to Te = 3P 4 Im {( iedq1 + ej \u03c0 6 iedq2 ) \u00b7 ( (Ldq1iedq1 +\u03c8edq1) + [ej \u03c0 6 (Ldq2iedq2)] )\u2217} = Te1 + Te2 + Te12, (35) where Te1 = 3P 4 Im { iedq1 \u00b7 (Ldq1iedq1 +\u03c8edq1)\u2217 } = 3P 4 ( \u03c8pmi e q1 + (Ld1 \u2212 Lq1)ied1i e q1 ) , (36) Te2 = 3P 4 Im { ej \u03c0 6 iedq2 \u00b7 (ej \u03c0 6 Ldq2iedq2 +\u03c8edq1)\u2217 } = 3P 4 ( \u03c8pm( \u221a 3 2 ieq2 + ied2 2 ) + (Ld2 \u2212 Lq2)ied2i e q2 + 3 \u221a 3 4 L\u03b4(i e2 q2 \u2212 ie2d2) ) , (37) Te12 = 3P 4 Im { iedq1 \u00b7 e\u2212j \u03c0 6 (Ldq2iedq2)\u2217 } + 3P 4 Im { ej \u03c0 6 iedq2 \u00b7 (Ldq1iedq1)\u2217 } = 3P 4 (1 2 (Ld1 \u2212 Ld2 \u2212 9 4 L\u03b4)i e d1i e d2 + 1 2 (Lq1 \u2212 Lq2 + 9 4 L\u03b4)i e q1i e q2 + \u221a 3 2 (Ld1 \u2212 Lq2 \u2212 3 4 L\u03b4)i e d1i e q2 + \u221a 3 2 (Ld2 \u2212 Lq1 \u2212 3 4 L\u03b4)i e q1i e d2 ) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001337_j.mechmachtheory.2017.12.026-Figure10-1.png",
        "caption": "Fig. 10. Determination of the carrier stiffness with aid of CAD software.",
        "texts": [
            " The variation of shared loads and contact stresses of an individual single tooth pairs, as well as the contact pattern at specific contact positions are analyzed and discussed later. The final stage of a three-stage planetary reducer (see Fig. 9 ) is chosen as the study case for analysis. The parameters of the planetary gear set are listed in Table 4 . Five planet gears are separated equally around the sun gear to carry out the high torque. The tooth profile is assumed as exact involute without any flank modification in the case. The equivalent stiffness of the carrier is obtained from FEM result by using CAD software Autodesk Inventor, see Fig. 10 . The total load with a value of 453.26 kN is applied either on the input side or on the output side of the carrier. The displacements \u03b4 and the calculated equivalent stiffness k of the carrier on the each side under various loading conditions are summarized in Table 5 . Similarly to FEM, the size of the discrete units affects also the analysis results. The suitable size is selected according to the convergence condition of the analysis results. The gearing data in Table 4 is used as example for the convergence test"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001875_20140313-3-in-3024.00016-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001875_20140313-3-in-3024.00016-Figure3-1.png",
        "caption": "Figure 3. Quadcopter configuration with Roll-Pitch-Yaw Euler angles [\u03c6, \u03b8, \u03c8]. [2]",
        "texts": [
            "  Gumstix: The QuaRC target computer. An embedded, Linux-based system with QuaRC runtime software installed.  Batteries: Two 3-cell, 2500 mA\u2219h Lithium-Polymer batteries.  Real-Time Control Software: The QuaRC-SIMULINK \u00ae configuration, as detailed in [4]. III. MATHEMATICAL MODEL OF THE QBALL-X4 The quadcopter considered in this work is an underactuated system with six outputs and four inputs and the states are highly coupled. Control of the quadcopter is attained by varying the speeds of each rotor. Fig. 3 shows the conceptual demonstration of a quadcopter. For the following discussion, the axes of the Qball-X4 are denoted as (x, y, z) and defined with respect to the configuration of the Qball-X4 as shown in Fig. 3. Roll, pitch, and yaw are defined as the angles of rotation about the x, y, and z axes, respectively. The global workspace axes are defined as (X, Y, Z) and defined with the same orientation as the Qball-X4 sitting upright on the ground. A non-linear model is proposed in [6]. For the purpose of simplicity, a linear model as furnished in [4] is adopted to design the controller. Each rotor produces a lift force and moment. The two pairs of rotors, i.e., rotors (\u201e1\u201f,\u201f2\u201f) and rotors (\u201e3\u201f,\u201f4\u201f) rotate in opposite directions so as to cancel the moment produced by the other pair"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001853_s10846-015-0233-z-Figure11-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001853_s10846-015-0233-z-Figure11-1.png",
        "caption": "Fig. 11 A schematic of a 3-DOF leg of a hexapod robot",
        "texts": [],
        "surrounding_texts": [
            "The following presents the experimental demonstration of foot force based reactive stability control using a radially symmetric hexapod robot. 4.1 Experimental Hardware Figure 7 shows the radially symmetric hexapod robot used in the experiment. The diameter of the platform is 300 mm. The legs of the robot have three joints as shown in Fig. 8. Each leg includes three separate segments which are connected together by revolute joints. The names, magnitudes, and limitations of the leg segments and joints are listed in Table 1. All of the legs are connected to the main body (platform) of the robot through hip joints. The hexapod robot consists of a Lynxmotion hexapod robot kit [25] and a Gumstix Verdex Pro XM4-BT COM tiny computer [26]. The hexapod robot has 18 HS-485HB servos controlled by a SSC-32 sequencer. There is a built-in proportional controller for each servo. Each leg consists of three servos and three leg segments. The Gumstix tiny computer acts as the high-level controller of the robot and communicates with the SSC-32 using UART via the Robostix expansion board. The Gumstix runs Linux 2.6 and is connected to the Netpro-VX expansion board for wireless connectivity. The housing for the Gumstix and expansion boards were fabricated using a rapid prototype machine. To physically measure the normal foot forces, the Lynxmotion hexapod robot, shown in Fig. 7, was equipped with Force-Sensitive Resistors (FSR-402) similar to [27]. The sensors were calibrated after they were embedded into the rapid prototyped housings, as shown in Fig. 9. The sensor data is read by the Robostix, shown in Fig. 7, and used for real-time monitoring of the stability. 4.2 Stability Metric A stability criterion needs to be defined within the controller to give the robot a sense of stability, to measure the stability of the robot, and to predict instability. Theoretically, any of the existing stability margins may be used for this purpose. However, the need for calculating the stability of the robot at any instant requires high frequency calculations within the controller which indicates a need for a concise stability margin with low computation cost and low sensor input information. The stability margin should also be able to present a quantitative stability extent which measures how close or far the robot is to the unstable or the maximum stable state. Hence, the FFSM [23], revisited in Section 3.2, is used since it only requires the measured normal foot/ground contact forces of the robot and provides a margin between zero and one which refer to the instability threshold and the maximum stability state of the robot, respectively. The conciseness and sensitivity of the FFSM make it efficient for use in an on-line and real-time controller to measure and predict the stability level of the system. 4.3 Inverse Kinematics The inverse kinematics is solved for both changing the location of the CG and changing the foot contact locations."
        ]
    },
    {
        "image_filename": "designv11_13_0003744_j.jbiomech.2018.12.024-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003744_j.jbiomech.2018.12.024-Figure1-1.png",
        "caption": "Fig. 1. Illustration of the collision model. A: The angle h is the angle between F and the vertical, k is the angle between V and the horizontal during the absorptive (left) and generative (right) phases of the stance. The collision angle / represents the deviation from an orthogonal relationship of F and V. B: When the average collision angle during the absorptive phase (U ) is greater than the average collision angle during the generative phase (U\u00fe), the collision is absorptive (left). When U <U\u00fe , the collision is generative (right) and when U = U\u00fe , the collision is \u2019pseudo-elastic\u2019 (middle).",
        "texts": [
            " According to these authors, one of the major sources of energy loss is the redirection of the motion of the center of mass of the body (COM), from downward to upward during stance. This down-to-up redirection is actively mediated by the legs and can, at first approximation, be considered as a collision with the ground. While a wheel prevents collision by maintaining a perpendicular relationship between the force and the velocity vector of its center of mass, a collision occurs if this relationship is not perpendicular; the angle of deviation from orthogonal relationship between the force vector (F) and the velocity vector (V) of the COM is called the collision angle / (Fig. 1A). Running results in a collision because F deviates from the vertical axis and V deviates from the horizontal axis in an opposite sense during stance (Fig. 1A). Based on a point mass collision model (described in details in the Methods section), Ruina et al. (2005) have shown that the collisional energy loss is minimized if the collision angle of the absorptive phase mirrors the subsequent collision angle of the generative phase. In this case, the collision is called a \u2018pseudo-elastic\u2019 collision and is energetically better than other ways of using legs. Only few studies have performed collision-based analysis on human gaits (Gutmann et al., 2013; Lee et al",
            " Because the down-to-up redirection of the COM is a key factor influencing the cost of running, we hypothesize that, at all speeds, the coordination strategies employed during this redirection will be close to a \u2018pseudo-elastic\u2019 collision around the \u2019optimal\u2019 slope of 6 . Ruina et al. (2005) proposed a point-mass collision model to evaluate the costs associated with the down-to-up redirection of the COM during stance. In this model, the runner is a point-mass (located at its COM) and the lower-limbs are assimilated to massless sticks that transmit forces along their axis. Three angles are defined (Fig. 1A): h is the angle between ground-interaction force F and the vertical, k is the angle between the velocity vector V of the COM and the horizontal and / the angle between V and a line orthogonal to F. Because the motion of the COM is a more forwards than up-downwards, the collision angles / are shallow and one can use small angle approximation: cos/ 1 and sin/ / (2005). Therefore, these last authors assume a conservation of linear momentum in the direction orthogonal to the leg and energy changes are accounted for by changes in the component of velocity along the axis of the leg between the beginning and the end of the stance (i",
            " The force signals were digitized by a 16-bit analog-to-digital converter at 1000 Hz. The horizontal (y) and the vertical (z) components of F and V were then computed using the methods described in details in Dewolf et al. (2016). The stance phase was defined as the period during which Fz was greater than 10 N. At each instant, the following parameters were calculated: the angle h between F and the vertical, the angle k between V and the horizontal and the collision angle / the deviation from an orthogonal relationship of F and V (Fig. 1A). These parameters were computed as in Lee et al. (2013): h \u00bc sin 1 F by=F \u00f05\u00de k \u00bc sin 1 V bz=V \u00f06\u00de / \u00bc sin 1 F V=FV\u00f0 \u00de \u00f07\u00de where by is a unit vector pointing horizontally forward and bz is a unit vector pointing vertically upward. Gutmann et al. (2013), Lee et al. (2013) computed the average of these angles throughout the stance phase, meaning that they did not consider the difference between landing and takeoff. Here, the stance phase was divided in two parts: the absorptive part during which / was negative and the generative part during which / was positive",
            " (2005) compare / at the initial and final instant of stance. At these moments, the horizontal and vertical GRF are rather small and the angle h is highly variable. Instead in this study, the absolute value of the angles h, k and / were averaged over the absorptive part (respectively H ;K ;U ) and over the generative part of the ground contact (respectively H\u00fe;K\u00fe;U \u00fe). The coefficient of generation eg was then evaluated from U\u00fe and U as: eg \u00bc U\u00fe U U\u00fe \u00feU : \u00f08\u00de Here, the collision is equivalent to a \u2018pseudo elastic\u2019 collision when eg \u00bc 0, i.e. U\u00fe \u00bc U (Fig. 1B). If eg \u00bc 1, the collision is exclusively generating, i.e. U \u00bc 0 and if eg \u00bc 1, the collision is a plastic collision exclusively absorbing, i.e. U\u00fe \u00bc 0. When running on a slope, the difference between the velocity in the direction orthogonal to the leg during the absorbing and generating phase was 4.5 \u00b1 2.3% (mean \u00b1 SD) of the average running speed. Because this result was similar to the difference obtained on the level (p = 0.236), we also assumed -as (2005)- that the energy change is completely accounted for by changes in the component of the velocity normal to the leg and we applied this pointmass collision model to our experimental data",
            " Interestingly, the model of the passive walker and runner of McGeer (1990) suggests that one could walk or run on a gentle downhill slope without any energy input, simply by taking advantage of gravity. In that case, the energy changes on a gentle negative slope balances the energy dissipated by joint friction, aerodynamic drag, etc. However, further studies should be considered to identify the reasons of the equality between the absorptive and generative collision angles on a 6 slope. The authors thank Prof L. Pen\u00e3ilillo for his contribution during the experiment and Mrs. Anne de Vinck for the drawings of Fig. 1. This study was funded by the Universit\u00e9 catholique de Louvain (Belgium), and the Fonds de la Recherche Scientifique (Belgium). Competing interests The authors declare no competing or financial interests. Braunstein, B., Arampatzis, A., Eysel, P., Br\u00fcggemann, G.P., 2010. Footwear affects the gearing at the ankle and knee joints during running. J. Biomech. 43, 2120\u2013 2125. Carrier, D.R., Heglund, N.C., Earls, K.D., 1994. Variable gearing during locomotion in the human musculoskeletal system. Science 265, 651\u2013653"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003283_0954406217745336-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003283_0954406217745336-Figure5-1.png",
        "caption": "Figure 5. The load and deformation of wheel\u2013rail system.",
        "texts": [
            "2 2 !2 1 \u00bc 2\u00f0 2!2 1!1\u00de !2 2 !2 1 8< : \u00f015\u00de where 1 and 2 are the damping ratios of two types of vibrations, according to experience 1\u00bc 0.05, 2\u00bc 0.01; !1 and !2 are the secondary natural frequency of the roller system, which can be determined by solving jK M!2 j \u00bc 0. Modeling of elastic shaft section 1. Shaft section division: The finite element analysis method was used here to simulate the flexible deformation characteristics of the shaft. The wheel set system\u2019s force analysis model is shown in Figure 5. The dividing principles were as follows: (1) the shaft shoulders was a division point to describe the different cross-sectional areas of the shaft section features; (2) the stress position of the shaft section was a division point used to describe the load characteristics of the shaft segment; (3) the shaft section division was made as uniform as possible. According to these principles, the rigid wheels were considered as a single particle on the axle, so the wheel\u2019s division position was considered the center of the rolling circle",
            " The normal contact force for left and right rails NLy and NRy can be defined using nonlinear contact theory21\u201323 NLy \u00bc 1 G dZL 3=2 NRy \u00bc 1 G dZR 3=2 ( \u00f019\u00de where G is the rail\u2019s contact constant; dZL and dZR are the relative normal elastic displacements of the left and right rails. For conical tread wheel, G fits the following G \u00bc 4:57R 0:149 10 8 \u00f020\u00de For wear type tread wheel, G fits the following G \u00bc 3:86R 0:115 10 8 \u00f021\u00de where R is the rolling radius of the wheel. The rail\u2019s relative normal elastic displacement was defined by left, right axle\u2019s flexible displacement and rail\u2019s displacement. The wheel and rail\u2019s force\u2013 deformation relationship is as Figure 5 shows. dZL \u00bc dwLc dwLr dZR \u00bc dwRc dwRr \u00f022\u00de where dwLc and dwRc are the displacements of the left and right axle at the wheel; dwLr and dwRr are the displacements of the left and right rail; and N0Ly and N0Ry and NLy andNRy are interactions. Consider the condition that the rail and wheel were separated, when the rail\u2019s normal forcewas 0, the left and right rails\u2019 normal forces NLy and NRy were defined using the equation below NLy \u00bc 1 G dZL 3=2 dZL 4 0 0 dZL40 ( NRy \u00bc 1 G dZR 3=2 dZR 4 0 0 dZR40 ( \u00f023\u00de When considering the rail\u2019s surface irregularity input Z(c,t), the rail\u2019s normal force fits the following NLy \u00bc 1 G dZL Zc Zt\u00bd 3=2 dZL 4 0 0 dZL40 ( NRy \u00bc 1 G dZR Zc Zt\u00bd 3=2 dZR 4 0 0 dZR40 ( \u00f024\u00de where Zc is the turnout impact model and Zt is the rail\u2019s irregularity input"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000780_j.jmapro.2015.11.009-Figure6-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000780_j.jmapro.2015.11.009-Figure6-1.png",
        "caption": "Fig. 6. Pinch effect.",
        "texts": [],
        "surrounding_texts": [
            "factur\n3\nI\nI\nw t g h b\nR\nw e E b i t h\nG\nw d a d h e\nm\nw\nx r s c l t e d t t u q x\n\u2212\nw\nx\nv\nj e\nj\narc between wires. In the case of current of 100 A and frequency of 20 Hz, distance for sustaining arc between wire tips is 4.2 mm. As the frequency increases to 200 Hz, distance for sustaining arc between wire tips monotonously decreases to 3.2 mm. Referring to\nS. Chen et al. / Journal of Manu\n. Results and discussion\nArc occurs when electric current passes through the twin wires. t can be governed by Ohm\u2019s law\n= V\nR (1)\nhere I, V, and R are, respectively, electric current, voltage and resisance. Eq. (1) implies that electric current easily passes through the ap between twin wires, or arc is easily ignited and sustained for a igh voltage or lower resistance between tips of twin wires given y\n=\nA (2)\nhere , and A are, respectively, gap between tips of twin wires, lectrical conductivity and an effective cross-sectional area of arc. q. (2) also indicates that the arc between tips of twin wires can e initiated earlier or easily sustained by decreasing distance, and ncreasing electrical conductivity and cross section of the arc. Elecron density N in arc increases with ionization produced by Joule eat given by\nJ = j2\n(3)\nhere GJ and j are, respectively, Joule heat and electric current ensity. Eqs. (2) and (3) indicate that electrical conductivity plays n important role in sustaining the arc between twin wires. To etermine electrical conductivity, Newton\u2019s second law of a simple armonic oscillator or the Lorentz model for dynamics of a bound lectron is used. That is,\nx\u0308 = eE \u2212 kx \u2212 m x\u0307 (4)\nhere displacement, acceleration and velocity of an electron are x,\n\u00a8 \u2261 d2x/dt 2\nand x\u0307 \u2261 v = dx/dt, respectively. Quantities e and m are, espectively, charge and mass of an electron, E electric field intenity, a measure of the rate of collisions per unit time, k the spring onstant with frequency \u03c90 = \u221a k/m, or k = m\u03c92\n0. The term on the eft hand side of Eq. (4) stands for inertia of the electron, whereas he first, second and last terms on the right hand side represent lectric field acting in the x-direction, a spring-like restoring force ue to the binding of the electron to the nucleus, and a frictionype force proportional to velocity of the electron, respectively. In he limit, frequency corresponding to the spring \u03c90 = 0 indicates an nbound electron in a good conductor. If electric current is of a freuency \u03c9, and expressed as E (t) = Eeiwt , and electron displacement (t) = xeiwt , Eq. (4) becomes\n\u03c92x + i\u03c9 x + \u03c92 0x = e\nm E (5)\nhere i \u2261 \u221a\u22121. The solution of Eq. (5) yields\n= eE/m\n\u03c92 0 \u2212 \u03c92 + i\u03c9\n(6)\nThe corresponding velocity of the electron is\n= i\u03c9x = i\u03c9eE/m\n\u03c92 0 \u2212 \u03c92 + i\u03c9\n(7)\nElectrical conductivity of a material is described by Ohm\u2019s law, = E. Electric current density j \u2261 Nev, where N is the number of lectrons per unit volume. It follows from Eq. (7) that\n= Nev = i\u03c9Ne2E/m\n\u03c92 0 \u2212 \u03c92 + i\u03c9\n\u2261 E (8)\ning Processes 21 (2016) 69\u201374 71\nElectrical conductivity, the reciprocal of electrical resistivity, is therefore dependent on frequency of electric current, given by\n(\u03c9) = i\u03c9\u03b50\u03c92 p\n\u03c92 0 \u2212 \u03c92 + i\u03c9\n(9)\nwhere \u03b50 is permittivity of free space, plasma frequency is\n\u03c9p = \u221a Ne2\nm\u03b50 (10)\nTo describe the arc or collisionless plasma \u03c90 = = 0, electrical conductivity given by Eq. (9) becomes pure imaginary:\n(\u03c9) = \u03b50\u03c92 p\ni\u03c9 (11)\nwhich indicates that electrical conductivity decreases or electrical resistivity increases with increasing current frequency and decreasing plasma frequency or electron density.\n3.1. The effects of electric current on the length of sustaining arc\nThe effects of magnitude and frequency of electric current on distance for sustaining arc between twin wires are shown in Fig. 3. It can be seen that distance for sustaining arc increases with electric current for a given frequency of electric current. For a given frequency of 20 Hz, for example, distance for sustaining arc between tips of twin wires increases from 4.2, 4.5 to 4.8 mm as electric current increases from 100, 120 to 140 A, respectively. Measurements also indicated that average welding voltage increases. Referring to Fig. 2(a) and (b), average voltage increases from 24.3 to 26.7 V. Similar trends can be seen for current frequencies of 50, 100, 150, and 200 Hz. Distance for sustaining the arc between wires increases as current frequency decreases. The decrease in electrical resistance is attributed to a decrease in electrical resistivity due to increased electron density or plasma frequency (see Eqs. (10) and (11)). Electron density increases with Joule heat (see Eq. (3)).",
            "72 S. Chen et al. / Journal of Manufacturing Processes 21 (2016) 69\u201374\nF b s\nF o c f i t d\ni t b i w w a w p i d p u i o f a i e c d\n3\nw w o 4 s i t c t\nig. 4. Distance for sustaining the arc between wire tips as a function of angle etween twin wires for a welding current of 120 A and frequency of 100 Hz of a quare wave.\nig. 2(a) and (c), average voltage also decreases. That is, in the case f current of 100 A average voltage decreases from 24.3 to 23.1 V as urrent frequency increases from 20 to 50 Hz. Similar trends can be ound for electric currents of 120 and 140 A. Eq. (11) shows that an ncrease in frequency of current through twin wires decreases elecrical conductivity. Increased electrical resistivity thus decreases istance for sustaining arc between twin wires.\n.4. The effects of wire radius on length of sustaining arc\nFig. 9 shows that in the case of current of 100 A an increase in ire radius decreases distance for sustaining the arc between twin ires. For a current frequency of 20 Hz, it can be seen that distance\nf sustaining the arc between twin wires decreases from 4.9 to .2 mm as wire diameter increases from 0.8 mm to 1.2 mm. Fig. 10 hows that average voltage for wire diameter of 0.8 mm is 28.5 V n comparison to 24.3 V for a wire diameter of 1.2 mm (Fig. 2(a)) in he case of current of 100 A and frequency of 20 Hz. Similar trends an be found for different current frequencies. Even though elecric resistance is inversely proportional to cross sectional area, a",
            "S. Chen et al. / Journal of Manufacturing Processes 21 (2016) 69\u201374 73\nd E c d\n4\n(\necrease in wire diameter increases Joule heat and electron density. lectrical resistivity increases as a result of a decrease in electric urrent density, leading to decreases in Joule heat and electron ensity.\n. Conclusions\nConclusions drawn are the following:\n1) Distance for sustaining the arc between twin-wire tips increases, as current frequency, angle between wires and wire diameter decrease, and average current magnitude increases.\n[\nngles between twin wires.\n(2) An increased distance for sustaining the arc between wire tips indicates a small electrical resistance between tips of wires, resulting in a low melting of wire tips, and a further low deposition rate. (3) A decrease in electrical resistance or increase in distance for sustaining the arc between twin-wire tips can be resulted from a decreased current frequency. This is because of a decrease in frequency-dependent electrical resistivity. (4) Distance for sustaining the arc between wire tips increases as angle between wires decreases. In view of the pinch effect, cross-sectional area decreases. An enhanced current density increases Joule heating and electron density, and decreases electrical resistivity and electrical resistance. Excluding electrical resistivity, the inversely proportional relationship between cross-sectional area and electrical resistance is comparatively minor. (5) Distance for sustaining the arc between wire tips increases as wire diameter decreases. Similar to item (4), increased electric current density enhances Joule heat, increases electron density, and decreases electrical resistivity. (6) Electrical resistance decreases as current magnitude increases. This is attributed to a decrease in electrical resistivity as a result of increases in Joule heat and electron density.\nAcknowledgements\nAuthors acknowledge the Program of Invited Distinguished Professorship (IDHT20150210) supported by Beijing Education Committee, P. R. China. This work is supported financially by National Natural Science Foundation of China (Grant nos. 51475009 and 51505009).\nReferences\n[1] Cui XM, Li LH, Zhang YH. Research situation of high efficiency welding process. New Technol New Process 2004;7:33\u20135. [2] Pan JL. The outlook of welding scientific research in the 21st century. In: The ninth welding conference of China. 1999. [3] Guan Q. The innovation and development of high energy beam welding/processing and solid connection technology. In: The 16th welding conference of China. 2011. [4] Shi CW, Zou Y, Zou ZD, Wu DT. Electromagnetic characteristics of twin-wire indirect arc welding. Chin J Mech. Eng 2015;28(1):123\u201331. [5] Ueyama T, Ohkawa T, Tanaka M. Affects of torch configuration and welding current on weld bead formation in high speed tandem pulsed gas metal arc welding of steel sheets. Sci Technol Weld Join 2005;10(6):750\u20139. [6] Tsushima S, Kitamura M. Occurrence of arc interaction in tandem pulsed gas metal arc welding. Sci Technol Weld Join 2007;12(6):523\u20139. [7] Harwig DD, Dierksheide J, Yapp D. Arc behavior and melting rate in the VPGMAW process. Weld J 2006;85(3):52\u201362. [8] Zhang YM, Jiang M, Lu W. Double electrodes improve GMAW heat input control.\nWeld J 2004;83(11):39\u201341.\n[9] Lu Y, Chen SJ, Shi Y, et al. Double-electrode arc welding process: principle, variants, control and developments. J Manuf Process 2014;16:93\u2013108. 10] Cao MQ, Zou ZD, Wang CM. Influence of welding current on arc characteristic of twin-wire indirect arc welding. Trans China Weld Inst 2005;26(12):47\u201350."
        ]
    },
    {
        "image_filename": "designv11_13_0000183_1077546314523030-Figure3-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000183_1077546314523030-Figure3-1.png",
        "caption": "Figure 3. Triangular element.",
        "texts": [
            " Equation (1) can then be simplified as @ @ x h3 @ p2 @ x \u00fe @ @ y h3 @ p2 @ y \u00fe Q i\u00bc x @ h p @ x \u00fe y @ h p @ y , \u00f03\u00de where i is the Kronecker delta, taken as 1 at the orifice, 0 for other parts; x and y are dimensionless bearing numbers; and Q is the flow rate factor for the gas mass flow into the orifice, referred to as the flow factor. For the bearings in steady-state operation, the first part on the right side of equation (1) is equal to zero. Equation (3) can then be used to calculate the gas consumption and pressure distribution of the aerostatic bearing. A triangular element is generally used to divide the gas film in the bearing. The shape of the element is shown in Figure 3 and the nodes in the elements are marked in a counter-clockwise order. The interpolation of the function f (z, x) must satisfy the following conditions: f (z, x) is continuous in an element and between any adjacent elements; it can depict equal conditions of gas film pressure within an element; and the gradient of the gas film pressure is continuous in the field of every element. The simplest interpolation function satisfying the above constraints is p \u00bc A\u00fe Bz\u00fe Cx \u00f04\u00de In the case of each node, the function of pressure is written as pi \u00bc A\u00fe Bzi \u00fe Cxi pj \u00bc A\u00fe Bzj \u00fe Cxj pm \u00bc A\u00fe Bzm \u00fe Cxm , \u00f05\u00de where zi and xi are the coordinates of node i, zj and xj are the coordinates of node j, and zm and xm are the coordinates of node m"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0003349_tac.2018.2817163-Figure1-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0003349_tac.2018.2817163-Figure1-1.png",
        "caption": "Fig. 1. This figure illustrates the first necessary condition. Given appropriate assumptions, the surface Lu\u2217V\u2217 (\u00b7) = 0 is tangent to the curve \u03c3n(M \u2212 R\u2217 \u2212 \u03b4M) = 0 for \u03b4M \u2208 S at M \u2212 R\u2217 \u2212 \u03b4M\u2217, where u\u2217 is the nth lsv and V\u2217 is the nth through mth rsv of M \u2212 R\u2217 \u2212 \u03b4M\u2217. For \u03b4M\u2217 to be a local minimum rank reducing perturbation for the property matrix R\u2217 the line connecting M \u2212 R\u2217 to M \u2212 R\u2217 \u2212 \u03b4M\u2217 must be perpendicular to the tangent surface Lu\u2217V\u2217 (\u00b7) = 0.",
        "texts": [
            ", \u2016\u03b4M\u2217\u2016F = r(M ;S,P) and rank(M \u2212 \u03b4M\u2217 \u2212 R\u2217) < n. We next provide intuitive development of the necessary conditions. Suppose that at the solution, rank(M \u2212 \u03b4M\u2217 \u2212R\u2217) = n\u2212 1. See Example 3 and the discussion following Assumption 3 for the case when rank(M \u2212 \u03b4M\u2217 \u2212R\u2217) < n\u2212 1. Let us first assume that the property matrix R\u2217 is fixed. For \u03b4M\u2217 to be the minimum norm rank-reducing perturbation for M \u2212R\u2217, the tangent plane to the hypersurface \u03a51 = {\u03b4M \u2208 S : \u03c3n(M\u2212R\u2217\u2212\u03b4M) = 0} must be perpendicular to the line connecting M\u2212R\u2217 and M\u2212R\u2217\u2212\u03b4M\u2217 (see Figure 1). For u\u2217 the nth left singular vector (lsv) and V\u2217 having columns equal to the nth through mth right singular vectors (rsv) of M\u2212R\u2217\u2212 \u03b4M\u2217, \u2016Lu\u2217V\u2217(M\u2212R\u2217\u2212\u03b4M\u2217)\u2016F = \u03c3n(M\u2212R\u2217\u2212\u03b4M\u2217) = 0. As will be seen in the proof of Theorem 1, the hyperplane \u03a52 = {\u03b4M \u2208 S : Lu\u2217V\u2217(M \u2212 R\u2217 \u2212 \u03b4M) = 0} is related to the tangent plane to \u03a51 at \u03b4M\u2217. Note, elements of \u03a52 are precisely the minima of \u2016Lu\u2217V\u2217S(\u03b4M)\u2212 Lu\u2217V\u2217(M \u2212R\u2217)\u2016F over \u03b4M \u2208 S. The minimum Frobenius norm rank reducing perturbation on M \u2212 R\u2217 is therefore the element of \u03a52 that has the least Frobenious norm",
            " Equivalently, (23) has the matrix form: H \u223c\u2032 (\u03b60, \u03c10) = [ \u2202H \u223c (\u03b6,\u03c10) \u2202\u03b6 \u2223\u2223 \u03b6=\u03b60 \u2223\u2223 \u2202H \u223c (\u03b60,\u03c1) \u2202\u03c1 \u2223\u2223 \u03c1=\u03c10 ] = \u2212 Re [ uHS1v, \u00b7 \u00b7 \u00b7 , uHSkv | uHP1v, \u00b7 \u00b7 \u00b7 , uHPrv ] = [LuvS(S1), \u00b7 \u00b7 \u00b7 | LuvP(P1), \u00b7 \u00b7 \u00b7 ], (24) Proof: See Appendix A. Thus Lemma 1 allows the replacement of the constraint H(\u03b6, \u03c1) = 0 in (8) by the Fre\u0301chet differentiable constraint H \u223c (\u03b6, \u03c1) = 0 when restricted to the set W . The next lemma proves that if the claimed necessary condi- tion 1a (or equivalently 1b) of Theorem 1 is not satisfied, then there exists a direction \u2206M \u2208 S to change the perturbation \u03b4M\u2217 on the tangent plane LuV (\u00b7) = 0 (see Figure 1). This new perturbation \u03b4M\u2217 + \u2206M \u223c may not be rank reducing, but will allow us to prove the existence of rank reducing perturbations with norms smaller than \u2016\u03b4M\u2217\u2016F . Lemma 2. Let M \u2208 Cn\u00d7m, \u03b4M0 \u2208 S , R0 \u2208 P satisfy rank[M \u2212 \u03b4M0 \u2212 R0] < n. Let u be the nth lsv and V have columns equal to the nth through mth rsv of M \u2212 \u03b4M0\u2212R0. Suppose LuV S surjective and \u03b4M0 6= (L\u2020uV S \u25e6LuV )(M\u2212R0). Then there exists a matrix \u2206M \u2208 S such that LuV S(\u2206M) = 0 (25) and \u3008\u03b4M0,\u2206M\u3009 < 0. (26) Proof: See Appendix A. Remark 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000599_j.engfailanal.2012.02.008-Figure18-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000599_j.engfailanal.2012.02.008-Figure18-1.png",
        "caption": "Fig. 18. The field of equivalent stresses on reinforced traverse beam.",
        "texts": [
            " With respect to supposed areas of material degradation and dimensional possibilities on the upper side of traverse beam it was proposed replacement of parts of upper flanges in the shape of circular sections with radius 200 mm according to Fig. 16. In Fig. 16 is given photograph of upper side of traverse beam after replacement of material in the corner of opening for bearing. In order to decrease stress levels in locations K and L it was recommended to reinforce traverse beams according to Fig. 17, where is given computational model for the finite element method. In Fig. 18 is shown the field of equivalent stresses on reinforced traverse beam for the loading by one full ladle. From comparison of equivalent stress fields on upper side of traverse beam before (Fig. 15) and after (Fig. 18) reinforcement results that reinforcement decreases in locations of stress concentrators K and L stress peaks by approximately 15%. On the basis of analysis obtained during solution problems of loading of casting pedestal the following conclusions can be stated. Due to periodical loading and unloading of wedge joints between traverse beams and middle part the successive drawing out of wedges occurred that results to irregular loading of traverse beam. This problem was solved with protection of upper wedges by stops"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0001991_demped.2015.7303696-Figure2-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0001991_demped.2015.7303696-Figure2-1.png",
        "caption": "Fig. 2. Two images of SKF 1205 ETN9 bearing.",
        "texts": [
            " It has been decided to evaluate the possible presence of the characteristic harmonic components derived from (5) and from the four previous equations, varying k from 1 to 5. The bearing faults have been simulated on a three-phase induction motor, rated power 2.2 kW, 2 poles, rated speed 2800 rpm, supplied by the mains at 400 V, 50 Hz. This motor is joined with a magnetic powder brake of 5 kW and 100 N\u22c5m through a flexible joint (Fig. 1). As in [13], the ball bearings SKF 1205 ETN9, doublecrown, self-aligning, with a plastic cage have been used, since they allow a quick and reliable disassembly of each rolling element (Fig. 2). The main bearing parameters are: outside diameter 52 mm, inside diameter 25 mm, 13 balls for each race (N = 26), D = 39.06 mm, d = 7.45 mm. The ball contact angle \u03b1 is estimated to be zero because of the self-alignment characteristic of the bearing. Generalized roughness was artificially created on all the balls of a bearing, at three progressive stages (Fig. 3): 1) step 1, early stage: the bearing was disassembled and all balls were degreased and immersed for 30 minutes in a solution composed of 5 ml of concentrated sulfuric acid (96%) in 50 ml of water"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0000368_aero.2013.6496903-Figure5-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0000368_aero.2013.6496903-Figure5-1.png",
        "caption": "Fig. 5. Calibration Stage",
        "texts": [
            " Only a coaxial cable is used between the capacitive gage and the probe. The shield of the cable is driven at the same voltage as the center conductor as done for triaxial cables. In this case, the ground connection between the probe and the gage must be made through a separate connection. Since one end of the measurement capacitance is the grounded turbine blade, this variation simplifies the cabling at the risk of introducing noise through the ground connection. A. Capacitance vs. Gap: the Need for Calibration Figure 5 shows the laboratory calibration x-y stage used to determine the relationship between the measured capacitance of a turbine blade and the gap between the blade and the capacitive probe. A CP-Ol Calibration Preamp is mounted on the y-axis stage with the capacitive probe flush with the surface of the front panel, which is similar to turbine installations. The blades are clamped on the x-axis stage, consisting of a Del-Tron crossed-roller slide and a Starrett digital micrometer. In operation, the blade is first centered over the probe and then moved with the micrometer",
            " Note that the guarded probe maintains a wider range over which the output tracks the ideal l/gap relationship. Note also that for systems not requiring a linearized output, the probe is potentially capable of operating well beyond the linearized maximum range with gaps on the order of 0.100\" in unguarded mode and 0.150\" in guarded mode. Beyond these extended limits, the capacitance drops below 0.001 pF, making accurate measurements problematic. Figure 7 shows the results with the same 0.075\" probe using a thin blade with dimensions of 0.010\" x I\". As with the thicker blade of Figure 5, the guarded configuration provides a higher capacitance and hence a higher SNR. Comparing Figures 5 and 6, the capacitance at small gaps is nearly a factor of ten higher for a blade of 0.250\" thickness compared to one of 0.010\". Note that even at small gaps, the capacitance for the thin blade does not follow the idealized l/gap relationship for either configuration. As a result, a linearized system that has been calibrated against a different target (such as an infmite surface) will exhibit significant errors when used with a thin blade",
            " A specially modified unit supplied to NASA Glenn Research Center achieved a bandwidth of 20 MHz. The HB-O 1 A Amplifier provides the buffered signal for guarded probes. C. Effect of Bandwidth on Accuracy This section describes a technique that can predict how the system bandwidth affects measurement accuracy, particularly for pulsatile measurements such as tip clearance of turbine blades. A transverse (y-axis) scan is initially generated for a particular blade at a particular gap (x-axis position) using the calibration stage of Figure 5. The horizontal axis is then mathematically converted to time based on blade tip velocity; the resulting \"pulse\" represents an ideal analog waveform that would be generated by a capacitive gage with infinite bandwidth. As examples, Figure 10 shows the normalized measurements for a guarded 0.075\" probe with a 0.010\" blade at gaps of 0.010\", 0.020\", and 0.050. Note that as the gap increases, the width of the pulse increases. For optical sensors, lateral spatial resolution is often measured by spot size, the width of the beam at which the intensity is 50% of the peak",
            " Stages in smaller engines generally have thinner blades and are monitored using smaller capacitive probes at smaller gaps. As a result, the blade pulse will generally be shorter, requiring even higher bandwidths for accurate measurements. The reverse occurs for larger diameter stages, which generally have thicker blades and would allow lower bandwidths for accurate measurement. However, an exception to this rule is a modern turbine blade design known as a squealer tip (2) (3). Figure 15 shows a blade model mounted on the calibration stage of Figure 5 and viewed from directly above the tip. The three tips are approximately 0.020\" in thickness separate by 0.060\"; the cavity between tips is approximately 0.060\" deep. Thus the total thickness of the blade at the tip is approximately 0.180\". The photograph was taken with the tip-to-blade gap set at 0.050\". Figure 16 shows the effect of bandwidth on the squealer model of Figure 15. Note that a bandwidth of 50 KHz virtually eliminates the detection of the individual squealer tips. If the system is calibrated using the fIrst tip, a 50 KHz bandwidth system eliminates the peak, potentially introducing signifIcant gap measurement errors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure10-1.png",
        "original_path": "designv11-13/openalex_figure/designv11_13_0002417_j.mechmachtheory.2015.08.022-Figure10-1.png",
        "caption": "Fig. 10. Influence of angle error of the Y2-axis on the contact region.",
        "texts": [
            " 9, it can be found that: (a) when the perpendicularity error of axis holes for mounting pinion and gear is positive, total contact regions will deflect to the topland of pinion and the root of gear in the conjugation process from the toe to the heel, respectively;(b) the influence of perpendicularity error in TL direction gradually reduces to zero, but the influence of axial error in TH direction gradually increases; and (c) the influence of perpendicularity error is less than that of linear errors. (5) Influence of angle error of the Y2-axis Fig. 10 shows the deflection tendency of contact region in TL and TH directions of pinion and gear when the angle error \u03b5y = 0.03 of Y2-axis. According to Fig. 10, it can be found that: (a) when the angle error of Y2-axis is positive, total contact regions deflect to the topland of pinion and the root of gear in the conjugation process from the toe to the heel, respectively; (b) the influence of angle error of the Y2-axis in TL direction gradually reduces to zero, but the influence of axial error in TH direction gradually increases; and (c) the influence of the angle error of Y2-axis is less than that of linear errors. By synthesizing Figs. 6\u201310, it can be concluded that among the linear error the influence of center distance error on the contact region is very obvious, and the influence laws of two components in the angle error on the contact region are similar; the influence of angle error of the Y2-axis is less than that of linear errors, which indicates that HNCGs is sensitive to the center distance error"
        ],
        "surrounding_texts": []
    }
]