[
    {
        "image_filename": "designv11_14_0000441_1.3662591-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000441_1.3662591-Figure1-1.png",
        "caption": "Fig. 1 Sketch of misaligned journal bearing",
        "texts": [
            ", October 20-22, 1959. NOTE: Statements and opinions advanced in papers are to be understood as individual expressions of their authors and not those of the Society. Manuscript received at ASME Headquarters, February 3, 1959. Paper No. 59\u2014Lub-3. misalignment. This compares favorably with similar-sized preloaded ball bearings for gyroscope rotors. Statement of Problem The problem is to calculate the pressure distribution and torque resulting from an angular misalignment of a gas-lubricated journal bearing. Fig. 1 shows a journal of radius r rotating at an angular velocity co within a stationary bearing of radius r + c and width (or length) b. The axes of the journal and bearing intersect at an angle S midway between the two ends of the bearing. In other words, the journal is not translated relative to the bearing but is merely rotated or misaligned. The bearing is operating in an atmosphere of gas at ambient pressure po. As the journal rotates it drags this gas through the gap h between the bearing and the journal, and sets up hydrodynamic (more properly aerodynamic or fluid dynamic) pressure forces on the journal and its bearing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003774_s0020-7403(96)00063-x-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003774_s0020-7403(96)00063-x-Figure7-1.png",
        "caption": "Fig. 7. Strain gage positions along the meridian profile: (a) final deformed shape; (b), (c) and (d) displacement pattern during the formation of the third fold.",
        "texts": [
            " Subsequently strain gages 16-20 are involved in the development of the third fold, strain gages 21-24 are involved in the initialization and development of the fourth fold, and finally strain gages 25 and 29 are involved in the development of the fourth fold. It is worth noting that the axial strains measured by strain gages 0, 1 and 2 are positive, due to the position on the external surface while the wall folds outward, and the axial strains measured by strain gages 3-7 are negative, due to the position on the external surface while the wall folds inward. This situation is schematically depicted in Fig. 7. It is also worth noting that the circumferential strains measured by strain gages 15 (placed at the same axial coordinate of the strain gage 0) and 24 (placed at the same axial coordinate of the strain gage 9) initially decrease until they become negative and then increase to positive high values. This can only be explained by an initial inward movement of the original cylindrical surface of the tube during the fold initialization phase, then followed by the outward motion of the surface during the fold formation phase, as illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002904_tmech.2021.3074800-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002904_tmech.2021.3074800-Figure7-1.png",
        "caption": "Fig. 7. A schematic view of a subject using the assistive walker during the STS transfer. hcom and vcom are respectively the height variation and velocity of the human trunk center of mass (COM), and h and v are respectively the height variation and velocity of the linear actuator.",
        "texts": [
            " \u2022 The STS transfer is a planar motion in the sagittal plane and bilaterally symmetric. Authorized licensed use limited to: Makerere University Library. Downloaded on May 17,2021 at 14:36:18 UTC from IEEE Xplore. Restrictions apply. 1083-4435 (c) 2021 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. \u2022 The height of the human chest is roughly equal to the one of the support plate during the STS transfer. Fig. 7 shows a schematic view of a subject using the assistive walker and performing the STS transfer. We use a human biomechanical model to represent the human body [6], [41]. In Fig. 7, the model composed by blue solid line represents the state at the beginning of the STS transfer, while the model composed by black dotted line represents the state during the STS transfer. Since the chest is assumed to mimic the trunk center of mass (COM) trajectory during the STS transfer [42], [43], we can have a simple approximate model between the linear actuator and the user: { hcom \u2248 kh vcom \u2248 kv, (2) where k is a constant determined by the lengths of the related assistive walker parts. Therefore, we can design the STS transfer controller by analyzing h and v of the linear actuator, instead of hcom and vcom of the user. The linear actuator represents the actual controlled object of the human-machine system, as shown in Fig. 7. During the STS transfer, the user would express his/her intention of standing up through the magnitude of the compression force that can be measured from the thin-film FSRs. The absolute value of the measured compression force reflects the intensity of users STS intention. Meanwhile, note that the direction of STS intention is opposite to the measured compression force. We define the intention force Fint that has the same absolute value but with opposite direction to the measured compression force"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001502_tia.2020.2972835-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001502_tia.2020.2972835-Figure1-1.png",
        "caption": "Fig. 1. Construction of a SAG mill with the stator and the rotor separately mounted.",
        "texts": [
            " Index Terms\u2014Alternators, ball milling, Fourier transform, gradient methods, pulse width modulation inverters, software algorithms, variable speed drives. I. INTRODUCTION THE production of cement requires machinery for crushing the cement clinker produced by a rotary kiln and to subsequently grind it to powder. The same procedure is followed in copper mining [1]. Crushed copper ore is ground to powder, from which the copper content is extracted in a chemical process. The production of fine powder is efficiently done in semiautonomous grinding (SAG) mills. Fig. 1 shows the construction of a SAG mill [2]. It consists of a rotating hollow cylinder of about 12 m in diameter. Two sets of bearings in ring-shaped structures support the cylinder to let it rotate. State-of-the-art gearless drives have about 76 salient poles attached around the mill cylinder. They form the rotor of a separately excited Manuscript received May 22, 2014; revised July 22, 2014, August 29, 2014, and October 19, 2014; accepted November 7, 2014. Date of publication December 9, 2014; date of current version May 8, 2015"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000728_icelmach.2016.7732503-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000728_icelmach.2016.7732503-Figure4-1.png",
        "caption": "Fig. 4. a) Flux distributions in the SynRM when the current density is 6 A/mm2. Position of the rotor in the direct axis. b) Flux distributions in the SynRM when the current density is 6 A/mm2. Position of the rotor in the quadrature axis.",
        "texts": [
            " This supposition is validated by comparing the variation of the inductances calculated by the FEA and the values obtained from experimental test (Fig. 5). The winding machine is assumed as a full-pitch, single layer inserted in a semiclosed slot. The coil conductor correspond of an enameled copper wire arranged in six parallel strands. For purposes of the simulation, it is simplified to a single-wire conductor with an equivalent conductor diameter and number of turns to match the inductance and resistance values calculated with the FEA and the experimental measurement. Fig. 4 shows the distribution of magnetic flux in the SynRM, predicted by the FEA, for the same value of stator current and magnetomotive force (MMF) but applied in the d-axis (a) and q-axis (b). Note that, as expected, the magnetic flux in the machine is considerably higher when the MMF stator acts on the direct axis (low reluctance) than for the case when it is applied to the quadrature axis (high reluctance). It also shows that, when the MMF is applied to the direct axis, the magnetic field lines cross the air gap and penetrate the rotor along the iron paths, but when the MMF acts in the quadrature axis, there is a small amount of magnetic flux that crosses the air gap and it closes through the saturated iron bridges and on the surface of the rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003766_0094-114x(95)00069-b-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003766_0094-114x(95)00069-b-Figure4-1.png",
        "caption": "Fig. 4. 6R kinematic chain. If all six joint angles are zero, the manipulator is in fully upward position, with all reference frames parallel.",
        "texts": [
            " Indeed, if joint i is prismatic, equations (10) and (11) simplify to Ji'b'= Ji\"e = ( O)i \u2022 (42) It is then straightforward to see that Pa and SA get some more zeros, but formally the relationships listed in Table 3 do not change. Hence also the symbolic expressions for the Jacobian derivatives do not change. 5. NUMERICAL RESULTS This section applies equation (36), for the joint angle derivative in the inertial representation, to the example of the 6R kinematic chain whose position and velocity kinematics are described by [18], see also Fig. 4. All joint angles equal to 1.0 rad result in the Jacobian MMT31/2--B 146 \u2022 H. Bruyninckx and J. de Schutter 0 -0 .54 --0.54 --0.77 0 - 0 . 8 4 - 0 . 8 4 0.49 y 1 0 0 --0.42 J = 0 841.47 1068.80 -718.62 0 -540 .30 -686.27 -1119.18 0 0 420.74 0 In this calculation, the length of the first link is 1000 mm, the third one 400 mm. Consider the derivative with respect S a ( J 3) = and, from (36), dJ ~q3 = SA(J3) J = -0 .5 9 - 0 . 6 4 -0 .27 -0 .43 0.77 - 0 . 6 4 617.28 206.94 (43) -142.34 -1124.12 423"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001385_tro.2019.2958207-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001385_tro.2019.2958207-Figure1-1.png",
        "caption": "Fig. 1. (a) TAILS: one of the fastest legged climbing robots. (b) Example of Full\u2013Goldman style dynamic climbing utilized by TAILS showing pendular COM trajectory and heading angle deviation (exaggerated for clarity).",
        "texts": [
            " These challenges include motion impediment versus foot placement constraints, postural changes (specifically enforced body shape changes), unusual (both holonomic and nonholonomic) motion constraints, and discrete state estimation, and control authority. The challenges described above, inherent to planning for a wide variety of legged locomotion, are demonstrated explicitly by dynamic vertical climbing, which imposes very unusual foot contact and motion constraints. Navigation within vertical climbing has been limited to quasi-static foot-step planning [14]\u2013[16], primarily because of the limited maneuverability of dynamic climbers. Recently, however, maneuverability has been demonstrated on the dynamic climbing platform TAILS, shown in Fig. 1, which is a member of the family of the fastest legged climbing robots (capable of climbing upward at over 1.5 body lengths per second). TAILS, using a single drive motor to actuate both front legs, is able to instantiate the animal inspired Full\u2013Goldman climbing dynamics [17]. The Full\u2013Goldman model captures the pendularlike oscillations [see Fig. 1(b)] characteristic of rapidly climbing animals. This style of climbing naturally provides passive self-stabilizing of heading angle using entirely feedforward control. Maneuverability 1552-3098 \u00a9 2019 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. was achieved via upward strafing and dynamic downward locomotion using the added rear leg actuation, which enables independent relative body roll and pitch control [18]",
            " These two examples highlight situations where considering the foot and the body separately (thinking about motion impediments versus foot placement constraints) could enable planning in more complex environments. In legged locomotion the propulsion mechanisms (the feet) are discretely placed and decentralized from a primarily passive body. The relative motion between these objects suggests that the feet and body can be treated as separate, but coupled, systems with distinct constraints. TAILS, shown in Fig. 1(a) [18], is a climbing robot, which demonstrates rapid (40 cm/s) ascension of vertical walls and instantiates the Full\u2013Goldman climbing dynamics. It does this by linearly retracting its forefeet through a sinusoidal extension trajectory locked 180\u25e6 out of phase using a single brushed dc motor. Directional adhesion [21] to the substrate is achieved through microspine arrays contained within the forefeet, which facilitate the rapid attachment and detachment necessary for dynamic climbing. The forefeet are connected to the body through a spring-loaded sliding wrist mechanism, which mitigates peaks in the wall reaction forces",
            " In previous results, it has been found that the body swing magnitude is much higher at low retraction frequencies and is attenuated at high frequencies, which causes the effective lateral size of the body to increase or decrease [18]. A specific example of this is the increased pitch required for strafing (nearly double vertical climbing), as shown in the video attachment. 4) State Estimation Challenges: During standard locomotion, the body uses pendular dynamic swings, which adds complexity to position tracking. While the average position change of a representative locomotion template shown in Fig. 1(b) is almost purely vertical, the instantaneous position of the robot\u2019s body deviates throughout the step from the prescribed linear path toward a goal. A perfectly well-behaved robot with known and consistent dynamics will have many options for types of motion planners as there is a clearly established relationship between the control space and the resulting motions. SBMPO was chosen as it is particularly adept at complex situations (such as the holonomic and nonholonomic constraints imposed by TAILS) by planning directly in the control space using a propagation relationship to map controls to whole-body velocities, which alleviates the need for an invertible model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000248_ecce.2019.8913075-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000248_ecce.2019.8913075-Figure4-1.png",
        "caption": "Fig. 4. 3-D FEA model of a DSSR AFIPM described in [21].",
        "texts": [
            " 100%s o g ECC g (5) For calculating the degree of eccentricity in case of combined eccentricity, the rotor is first deviated with required \u03b2 to achieve necessary SEF and then horizontal shifting is done for a particular ECC. IV. 3-D FEA MODELING OF DSSR AFIPM MOTOR Among the various techniques of eccentricity modeling in AFPMs, 3D-FEA model can precisely demonstrate the impact of the eccentricity because it can accurately measure the air gap region [15]. To model the static eccentricities in FE analysis, a DSSR AFIPM with V-shaped permanent magnet described in [21] has been used in this study. The DSSR AFIPM has higher inductance and saliency, which enables it to have a wider speed range. Fig. 4 shows the FEA model of the V-shaped DSSR AFIPM. Table I shows the design specifications of the motor. It has 12 slot 10 pole structure with concentrated winding scheme. The magnets used in the motor is NdFeB35. V. SIMULATION RESULTS 3D-FEA has been carried out for all three types of static eccentricity in DSSR AFIPM. The degree of eccentricity has been varied from 0% to a maximum value of 40%. To generate static angle eccentricity, rotor has been axially deviated along y-axis. Based on the design specifications of the DSSR AFIPM, deviation angle for different degrees of SEF has been calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000918_iet-epa.2018.5918-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000918_iet-epa.2018.5918-Figure1-1.png",
        "caption": "Fig. 1 In six-phase system (a) D3VSIs fed 6PIM, (b) Voltage space vector in both planes",
        "texts": [
            " In this paper, total 18 different SVPWM sequences are discussed: among which some are implemented in various papers while 10 new techniques are introduced in this paper. Total six PWM techniques are selected for comparative analysis. A comparative study, corresponding to three continuous and three discontinuous SVPWM techniques, based on HDF, total harmonic distortion (THD), torque ripple and switching losses are discussed. The SPIM has two sets of stator windings separated by 30\u00b0 as shown in Fig. 1a. The dual D3VSIs produce six-phase voltages which are transformed by applying Clarke's decoupling transformation matrix and are defined as (see (1)) . The D3VSIs produce six-phase voltages which are switched to d\u2013q and x\u2013y planes using (1) as shown in Fig. 1b. The third subspace 01\u201302 is excluded due to isolated neutral points of two three-phase windings. In Fig. 1b inverter output voltage vectors are represented as the decimal number of binary switching pattern (Sc2, Sb2, Sa2, Sc1, Sb1, Sa1) corresponding to the state of the upper switch of respective inverter leg. All state voltage vectors are categorised based on their magnitude (VL: large, VML: mediumlarge, VM: medium, VS: small and V0: zero) as shown in Fig. 1b. IET Electr. Power Appl., 2019, Vol. 13 Iss. 11, pp. 1753-1762 \u00a9 The Institution of Engineering and Technology 2019 1753 Small vectors of the d\u2013q plane are mapped into large vectors in the x\u2013y plane and never utilised in SVPWM techniques, thus removed in Fig. 1b. The sinusoidal PWM (SPWM) and double zero sequence injection PWM (ZSPWM) utilise five active vectors in each sector and are represented as 24-sector PWM techniques. For further reference, switching sequences utilised by carrier-based PWM techniques are well described in [10]. Table 1 summarises the switching sequence used in sector-1 and sector-3. A further resultant sequence in sector-1 has been derived by transforming the sequence of sector-3 into sector-1 so as obtain trajectories are similar to the sequence of sector-1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000067_metroi4.2019.8792876-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000067_metroi4.2019.8792876-Figure5-1.png",
        "caption": "Fig. 5. The tolerances imposed to the component during the design phase.",
        "texts": [
            " To this end, the characterization of the dimensions, and thus the distortions, was conducted by implementing two series of measurements namely one after the SLM process and one after the thermal treatment. Consequently, the comparison of the two series of measurements may identify the corresponding deviation from the nominal design and dimensions. Considering the above, the measurements were taken from the upper surface of the previously described component of the satellite due the fact that it should host the closing belt of the satellite antenna. The tolerances described by the designer are 0.05 mm as seen in Fig. 5. Consequently, the measurements\u2019 campaign starts with the identification of the component within the overall measuring volume as seen in Fig.6, thus to set the local-global coordinate system. To this end, the holes of the supporting screws utilized as their nominal coordinates were already known. The centers of these holes were identified easily and are consequently considered as the centers of the potential local Cartesian coordinate systems BF1, BF2 and BF3 respectively as seen in Fig 7a. After having identified the coordinates of the points BF1, BF2 and BF3 respectively, the global coordinate system starting from the point BF1 was considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001218_s42417-019-00182-5-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001218_s42417-019-00182-5-Figure1-1.png",
        "caption": "Fig. 1 A 3D model of the spindle feed system",
        "texts": [
            " Meanwhile, the Z-axis motor generates heat, from the friction of the bearings at the front and rear ends of the ball screw, friction between the ball screw and the transmission nut, and friction between the rolling guide and the slider. By definition, all friction actions also generate the heat. However, heat from cutting is also transmitted to the Z-axis through the spindle components which cannot be ignored. All the above heat sources affect the position accuracy of the feed motion. An example of this type of mechanism is shown in Fig.\u00a01. Q1\u2013Q6 denote the heat sources of the spindle feed system. 1 3 The spindle feeding system is constantly affected by the cutting force, cutting heat and friction heat during the machining process. Over a small range, these effects produce a high stress, strain, strain rate and instantaneous temperature rise, which cause structural deformations. This process is representative of typical thermo-mechanical coupling as opposed to the superposition of separate deformations. If the cutting force deformation is u1 , the thermal deformation is u2 and the thermo-mechanical coupling deformation is u , where u1 + u2 \u2260 u ",
            " Figures\u00a05 and 6 show the deformations of the upper and lower ends of the screw due to the milling forces. These figures indicate that the deformations in the Z-direction at both the upper and lower ends are relatively small, while the deformations in the X- and Y-directions are is relatively large, with a maximum of approximately 1.65\u00a0\u03bcm. The spindle feed system and the associated components are sensitive to temperature changes that could deform the mechanical structure. The heat sources (Q1\u2013Q6) of the machine tool spindle feed system are shown in Fig.\u00a01. Among them, the temperature values for Q1\u2013Q5 are obtained by from the heat model in Eq.\u00a0(6). Based on the heat milling thermal theory model, the temperature of the cutting edge reaches a steady state after approximately 0.6\u00a0s during high-speed milling. The results show that the milling temperature at the cutting edge fluctuates around 50\u00a0\u00b0C. Therefore, this value is applied as a constant at the edge of the milling cutter. The temperatures applied to other heat sources of the screw system shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001184_j.ijfatigue.2019.105281-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001184_j.ijfatigue.2019.105281-Figure11-1.png",
        "caption": "Fig. 11. FE model and boundary conditions (crack length a=1mm).",
        "texts": [
            " As \u0394KII reaches the threshold value \u0394KII,th, the mode II fatigue crack starts to branch in the direction of mode I growth. The branch angle is approximately\u00b1 70.5\u00b0, where the tangential normal stress reaches its maximum value [32]. A gap appears between the crack surfaces because of the material loss caused by wear. Four specimens were tested in this study and the mode II fatigue crack length was determined to be 1.01, 1.18, 0.96, and 1.08mm, respectively. To determine \u0394KII,th of the test wheel steel, the FE code ABAQUS was used to calculate \u0394KII for different crack lengths. The FE mode shown in Fig. 11 was created based on the shape and dimensions of the test specimen. The specimen was modelled as elastic material with a Young\u2019s modulus of 210 GPa and Poisson\u2019s ratio of 0.3. The crack length a was set to 0.5, 0.8, 1.0, 1.2, and 1.5 mm, respectively. To improve the analysis accuracy, the mesh size around the crack tip was refined to be 0.4 \u03bcm. The penalty approach was used to impose frictional constraints on the crack surfaces. The friction coefficient between the crack surfaces was selected to be 0.6, which is the typical value for a friction pair composed of ferrite\u2013pearlite steels [33]. As shown in Fig. 11, a symmetry constraint (U3=UR1=UR2=0) was applied on the surface S1 in the z-direction. Fixed constraints (U1=U2=U3=0) were applied on the surfaces contacting the jig. An S value of 10.58 kN was applied to points A and B to simulate the clamp force generated by the bolts. The test load P=5.9 kN was then applied to point A. The stress intensity factors were directly extracted from ABAQUS, which makes use of the contour integral method to compute the stress intensity factors. Fig. 12 shows the stress state of the region near the neutral section under loading conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000506_s11771-016-3101-5-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000506_s11771-016-3101-5-Figure2-1.png",
        "caption": "Fig. 2 Coordinate systems applied for gear generation",
        "texts": [
            "2 Part (b) of generating surface for gear circular profile head-cutter In the coordinate system S0, surface (b) g\u03a3 of the head-cutter is represented by vector function (b) 02r as follows: ( ) ( ) ( ) ( ) ( ) w2 2 2 2 b 2 2 w2 2 2 202 2 2 sin cos , sin sin 1 cos X X \u03c1 \u03bb \u03b8 \u03bb \u03b8 \u03c1 \u03bb \u03b8 \u03c1 \u03bb    =    \u2212 \u2212   r , 2 2 \u03c00 2 \u03bb \u03b1\u2264 \u2264 \u2212 (5) where ( )w2 2 2 2 2sec tanX R \u03c1 \u03b1 \u03b1= \u2212 \u00b1 \u2212 (6) According to Eq. (3), the unit normal vector of the gear generating surface (b) g\u03a3 is determined as ( ) ( ) 2 2 b 2 2 2 202 2 sin cos , sin sin cos \u03bb \u03b8 \u03bb \u03b8 \u03bb \u03b8 \u03bb    =      n (7) The upper and lower signs in Eqs. (5), (6) and (7) correspond to generation of the convex and concave sides of the gear tooth surface, respectively. 2.1.3 Equations of gear tooth surface Coordinate system Sm{Xm, Ym, Zm} is rigidly connected to the cutting machine (Fig. 2). The top, bottom and upper right of Fig. 2 are the machine front view, the machine bottom view and the side view (the projection of head cutter), respectively. The cradle is rotated about the Ym-axis; the p2-axis is projection of the gear axis in the XmOmYm-plane. The points M, P and O0 are the reference point of the tooth surface, the mean contact point and the center of the head cutter, respectively. O2 is the cross point of gear, and Om is the machine center. Some of the machine-tool settings of the given hypoid generator are: the machine root angle of gear \u03b3m2, the machine center to back XG, the cradle angle of gear q2, the radial distance of gear Sr2, the blank offset Em2, the sliding base XB2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002221_0278364919897134-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002221_0278364919897134-Figure15-1.png",
        "caption": "Fig. 15. Buoyancy test cylinder with fiber-reinforced membrane as a buoyancy bladder. Pumping water in and out of the bladder changes the density of the device as a whole, causing it to ascend or descend in the water column.",
        "texts": [
            " Both membranes were also inflated to the same hemispherical shape before being loaded by the same aluminum block to directly compare their shape change. 4.3.2. Internal loading. An internal load was exerted on the membrane tangent to the clamp base by inflating the membrane with water. Images were taken at various pressures to compare the loaded membranes\u2019 geometry. To demonstrate this fiber-reinforcement technique\u2019s utility in a practical application, we implemented it on our group\u2019s submersible buoyancy test cylinder, shown in Figure 15. Previously, we filled or emptied an inextensible plastic swim bladder with water to change the density of the test cylinder and cause it to ascend or descend in the water column (Sholl and Mohseni, 2019). This technique has several disadvantages, including an inability to control the shape of the bladder, which can kink itself or push into other components, and an inability measure the volume of the bladder, which maintains the same pressure as the internal cavity of the test cylinder throughout operation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure5-1.png",
        "caption": "Fig. 5. Rotation with b = p=6.",
        "texts": [],
        "surrounding_texts": [
            "The objective of our method is to estimate all of the unique components of I and rC=B. The strategy relies on being able to translate the object by rB and then rotate it by R with respect to frame O which is defined by the apparatus. The term IR in Eq. (15) can be expanded to: IR \u00bc R Ixx \u00fem yC 2 \u00fe z2C Ixy mxCyC Ixz mxCzC Ixy mxCyC Iyy \u00fem xC2 \u00fe z2C Iyz myCzC Ixz mxCzC Iyz myCzC Izz \u00fem xC2 \u00fe y2C 0 B@ 1 CART \u00f016\u00de The selection of R is very important. It was found that the rotations presented in Figs. 5 and 6 are of mathematical and experimental convenience. For experimental testing, the object is first rotated about yO by angle b, then about x1 by angle a, as shown in Figures 5 and 6, where the bases x1; y1; z1 and x2; y2; z2 are the result of the first and second rotations respectively. The resulting rotation matrix is written as: R \u00bc Ry;bRx;a \u00bc cos b\u00f0 \u00de sin a\u00f0 \u00de sin b\u00f0 \u00de cos a\u00f0 \u00de sin b\u00f0 \u00de 0 cos a\u00f0 \u00de sin a\u00f0 \u00de sin b\u00f0 \u00de cos b\u00f0 \u00de sin a\u00f0 \u00de cos a\u00f0 \u00de cos b\u00f0 \u00de 0 B@ 1 CA \u00f017\u00de We focus on the bottom-right component of IR(15), denoted as IRzz, whose value can be measured directly (c.f. Section 4): i 1 2 3 IRzz \u00bc Ixx sin 2 b\u00f0 \u00de 2Ixy sin a\u00f0 \u00de sin b\u00f0 \u00de cos b\u00f0 \u00de \u00fe 2 cos a\u00f0 \u00de cos b\u00f0 \u00de sin b\u00f0 \u00de mxCzC Ixz\u00f0 \u00de\u00f0 \u00fe sin a\u00f0 \u00de cos b\u00f0 \u00de Iyz myCzC \u00de \u00fe sin2 a\u00f0 \u00de cos2 b\u00f0 \u00de Iyy \u00fem xC2 \u00fe z2C \u00fe cos2 a\u00f0 \u00de cos2 b\u00f0 \u00de Izz \u00fem xC2 \u00fe y2C \u00femxCyC sin a\u00f0 \u00de sin 2b\u00f0 \u00de \u00femy2C sin 2 b\u00f0 \u00de \u00femz2C sin 2 b\u00f0 \u00de \u00f018\u00de where from (13), xC \u00bc xB \u00fe xC=B \u00f019\u00de yC \u00bc yB \u00fe yC=B \u00f020\u00de zC \u00bc zB \u00fe zC=B \u00f021\u00de The choice of rotation and translation configuration parameters b;a; xB; yB, and zB can simplify the experimental determination of the desired inertia tensor I and CM location vector rC=B. This method requires a minimum of nine experimental configurations to estimate nine unknowns: Ixx; Iyy; Izz; Ixy; Ixz; Iyz; xC=B; yC=B; zC=B. It is assumed that the mass of the object, m, is known or can be measured. Motion measurements allow estimation of IOzz, the total mass moment of inertia about the oscillation axis, which is the summation of the inertia of the rotating assembly (plate, attachment blocks, etc) IP and the inertia of the measured object IRzz. For a given experiment with index i, the inertia corresponding to the measured object can be found from: IRzz;i \u00bc IOzz;i IP;i :\u00bc Ie;i \u00f022\u00de The terms IP;i can be estimated through direct experiments, which will be done in Section 4.7. It is worth noting that IOzz;i is assumed to be constant during a given experiment, which means the parameters b;a; xB; yB, and zB are configured before the tests, and remain constant."
        ]
    },
    {
        "image_filename": "designv11_14_0002046_icem49940.2020.9270988-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002046_icem49940.2020.9270988-Figure9-1.png",
        "caption": "Fig. 9 Modal shapes of several structural modes",
        "texts": [
            " 2) Uneven pole magnetization Uneven magnetization introduced new stress harmonics as sidebands of existing harmonics in even magnetization condition (cf. Fig. 8). Therefore, all mechanical orders and wavenumbers exist in the 2D FFT of radial magnetic stress. Same harmonic content is found for tangential stress. 1) Main structural modes Main stator structural modes are obtained from numerical modal analysis and presented in Table 3. The main mode of interest is the breathing mode around 5450 Hz (cf. Fig. 9) which is likely to be excited by radial pulsating forces. Other radial modes (m = {1,2,3,4}, n = 0) can be excited in uneven magnetization condition since all wavenumbers exist in magnetic radial stress harmonic content. Besides, both tooth rocking modes at 4513 and 4519 Hz are likely to be excited by tangential stress harmonics of wavenumber r = \u00b18. TABLE 3 ORDERS AND NATURAL FREQUENCIES OF STRUCTURAL MODES Mode identification (m,n) (m: circumferential order; n: longitudinal order) Natural frequency (Hz) Yoke \u201cbreathing\u201d mode (0,0) 5450 Yoke \u2018bending modes (1,0) 963 & 970 Yoke \u201covalization\u201d modes (2,0) 995 & 1000 Radial modes (3,0) 1732 & 1737 Radial modes (4,0) 2698 & 2703 Tooth rocking modes (8,0) 4513 & 4519 2) Displacement FRF, r = 0,2,8 RMS value of displacement FRF under radial (respectively tangential) unit-magnitude force waves of wavenumber r = 0,2,8 are illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002329_042033-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002329_042033-Figure2-1.png",
        "caption": "Figure 2. Simulation models: (a) \u2013 attachment; (b) \u2013 wheels.",
        "texts": [
            " By adjusting the position of the ball, the center of mass of the simplified model is combined with the real operational center of mass of the tractor. Further, for the layout of the simulation model of the tractor, simplified models of the front and rear three-point linkage have been created. They have a simple parameterized geometry and retain mass-inertial parameters. This allows you to use models with the least load on the computer and the minimum probability of errors. For simulation studies, the three-point linkage model is supplemented by virtual elastic elements (figure 2, a). For this, a central bracket is additionally created synchronously moving in a longitudinally vertical plane with lower draft link. It is necessary for mounting virtual springs 1. The second mounting point is located on the lower draft link. Another pair of virtual springs 2 is installed at the central points of attachment of the hydraulic cylinders. By specifying the stiffness of the springs and damping coefficients, a wide range of operating conditions can be modeled. The lower draft links are moved due to linear virtual engines 3 applied in parallel to the axis of the cylinder. To simulate the lateral displacement of the tractor resulting from tire deformation, the wheel model is also supplemented with elastic elements (figure 2, b). ICMSIT 2020 Journal of Physics: Conference Series 1515 (2020) 042033 IOP Publishing doi:10.1088/1742-6596/1515/4/042033 A virtual spring 4 connects the wheel disk and the tire having the possibility of axial displacement without relative rotation. By selecting the elastic coefficients and damping of the virtual spring, tire deformation can be simulated. To study the lateral stability, the machine-tractor unit is installed on a virtual stand (figure 3). It consists of a fixed base and a platform that changes the angle of inclination"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000844_1464419318819332-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000844_1464419318819332-Figure3-1.png",
        "caption": "Figure 3. Combined IRD\u2013ORD\u2013RD interaction.",
        "texts": [
            " The defect amplitude Bdi as a result of the inner race defect to roller interaction was described by Patel and Upadhyay18 as Contribution of ORD. The defect amplitude, Bdo as a result of the outer race defect to roller interaction was described by Patel and Upadhyay18 as where \" is the initial angle of defect. Coupled IRD\u2013ORD\u2013RD. For IR and OR interaction with RD, Bd \u00bc Bdri and Bd \u00bc Bdro respectively. For IRD and ORD, Bd \u00bc Bdi and Bd \u00bc Bdo respectively. For combined IRD\u2013ORD\u2013RD interaction as per Figure 3, Bd \u00bc Bdri \u00fe Bdro \u00fe Bdi \u00fe Bdo. A model with nonlinear stiffness and damping is prepared here with fixed OR. However, nonlinearity due to the defect is foremost. A non-Hertzian model is applied for the contact stiffness formulation. Load\u2013 deflection relation was described by Lundberg and Palmgren41 as \u00bc 3:84 10 5 A0:9 g0:8 \u00f08\u00de The deformation between roller and races are due to radial load, roller tilt and skew, radial gap and fault on the roller or united fault on IR-OR-R. Deflection of IR 2pl and OR 1pl interaction with roller is denoted by 1pl \u00bc 1p b l 1 2 p \u00fe q 2 l 1 2 q\u00fe 1\u00f0 \u00de b dm \u00fe d \"2p Z 2 Bd \u00f09\u00de Bdo \u00bc r r cos J 2r sin o J mod di, 2 \u00f0 \u00de \"\u00f0 \u00de , 04mod di, 2 \u00f0 \u00de \"5 d1 r r cos J 2r , d14mod di, 2 \u00f0 \u00de \"5 d d1 r r cos J 2r sin o J mod di, 2 \u00f0 \u00de \"\u00f0 \u00de , d d14mod di, 2 \u00f0 \u00de \"5 d 8>< >: \u00f07\u00de Bdi \u00bc r r cos J 2r : sin i J mod di, 2 \u00f0 \u00de\u00f0 \u00de , 04mod di, 2 \u00f0 \u00de5 d1 r r cos J 2r , d14mod di, 2 \u00f0 \u00de5 d d1 r r cos J 2r : sin i J mod di, 2 \u00f0 \u00de\u00f0 \u00de , d d14mod di, 2 \u00f0 \u00de5 d 8>< >: \u00f06\u00de 2pl \u00bc 2p \u00fe b l 1 2 p \u00fe l 1 2 q\u00fe 1\u00f0 \u00de b 2 dm \u00fe d \"2p Z 2 Bd \u00f010\u00de where Bd is an amplitude either due to RD or due to combined IRD\u2013ORD\u2013RD"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002037_s00170-020-06366-8-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002037_s00170-020-06366-8-Figure3-1.png",
        "caption": "Fig. 3 Mould inserts configuration with new CCC design",
        "texts": [
            " Mechanical properties and microstructure of the hybrid-build parts were also evaluated for proof of practicality. The results could provide a cost-effective option for mould makers and plastic products manufacturer in choosing new AM technologies for their applications. Four mould inserts, which are part of an existing injection mould for a plastic round pot, were chosen for the study. The cooling circuits of these four inserts were redesigned as conformal cooling channels using the Siemens NX CAD software package [14], with CCC design criteria based on current knowledge [9, 15]. Figure 3 shows the selected inserts in the mould, with redesigned configurations. To illustrate the MSHAM design concept, Fig. 4 shows the details of the redesigned Wedge Insert \u201cA\u201d. The design parameters for the conformal cooling channels were: \u2022 Channel cross-sectional shape: circular \u2022 Channel size: 2 mm and 4 mm diameter \u2022 Channel wall to cavity wall distance: 2.4 mm, 3.8 mm and 4.3 mm For availability reason, the powder material chosen for the AM process in this study was AlSi10Mg aluminium, with the chemical composition shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure23-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure23-1.png",
        "caption": "Fig. 23. Location of UAV centre reference point. Frame O origin is located under and forward.",
        "texts": [
            " A misalignment would introduce systematic errors into the estimates. Employing the procedure outlined in Section 4.6, the estimated inertia and position of the centre of mass were estimated to be: 2. The UAV supported by a lightweight metal structure and gimbal mounted on torsional platform. Two configurations (a) and (b) are shown. 6 rations for UAV experiments. IUAV \u00bc 0:00210 0:00023 0:00018 0:00023 0:00217 0:00019 0:00018 0:00019 0:00336 0 B@ 1 CA kg m2 \u00f086\u00de rC=B UAV \u00bc 0:0065 0:0000 0:0678\u00f0 \u00deT m \u00f087\u00de As shown in Fig. 23, we define a reference point on the UAV located at the geometric centre of the four propellers and on the surface of the battery attachment plane. By direct measurement, the position vector of the point relative to the gimbalcentered frame O is rp \u00bc 6;0;67:9\u00f0 \u00deT mm. Combined with the rC=B UAV estimate, the position of the centre of mass of the UAV with respect to the geometric reference point is thus rCM UAV \u00bc 0:5 0:0 0:1\u00f0 \u00deT mm \u00f088\u00de The small values in rCM UAV are as expected, since UAV designers place the centre of mass as close as possible to the centre of pressure in order to simplify flight control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure12.3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure12.3-1.png",
        "caption": "Fig. 12.3 a General working diagram of the aerial manipulation system. b Geometric parameters of manipulator",
        "texts": [
            " In our study, a robotic arm is proposed with the first two links of that robotic arm plays a role as counterbalance weight instead of using another end-effector is a counterbalance weight like in [14]. In Fig. 12.2 shown that with the pose of the robotic arm in Fig. 12.2a, the COM of the robotic arm is kept closer to vertical axis of UAV. In Fig. 12.2b, the COM is quite far from the vertical axis of UAV and we will not consider the pose of the robotic arm as stable. We offer the general working diagram of the aerial manipulation system presented in Fig. 12.3a: S1 is a UAV, S2 is the space where the robotic arm is placed when it is in the non-working state (parking and takeoff stages), S3 is the workspace of robotic arm. S3 is the space containing objects to manipulate. In this study, we will analyze the theta deflection angle between the links of the arm (Fig. 12.3b) when the robotic arm moves from S2 to S3. Through that we will find out the optimal theta angle set, so that the robotic arm moves but still keeps its the change of COM is minimal along the horizontal axis X. The objective is always to keep the COM of the robotic arm moving on the close space of vertical axis Y. Suppose that there is a manipulator consists of five links mounted on UAV. The kind of joint between two consecutive links is only revolute joints such as Fig. 12.3b. So we have geometric parameters table of the serial chain manipulator in Table 12.1. The purpose of this forward kinematic is determining the coordinates of the joints Ji (i = 1 \u00f7 5) so that we can calculate the COM coordinates of the manipulator. Pay attention that J5 is like the end-effector, the coordinates of J0 are (x0 = 0; y0 = 0). ai is the distance from zi\u22121 to zi along xi\u22121, \u03b1i is the angle from zi\u22121 to zi about xi\u22121, where ci = cos \u03b8i ; si = sin \u03b8i ; si j = sin(\u03b8i + \u03b8 j ) = cos \u03b8i sin \u03b8 j + cos \u03b8 j sin \u03b8i ; ci j = cos(\u03b8i + \u03b8 j ) = cos \u03b8i cos \u03b8 j + sin \u03b8 i sin \u03b8 j ; Coordinate frame i can be located relative to coordinate frame i \u2212 1 by executing a rotation through an angle",
            "5, where: 1\u2014the position of manipulator\u2019 end-effector when the system is in parking and takeoff stages; 2, 3, 4\u2014the position of manipulators\u2019 end-effector when the system moves to point to the workspace; 5\u2014the position of manipulators\u2019 end-effector when the system is at workspace. The end-effector has Fig. 12.4 Working diagram of aerial manipulation system Parking stage Take-off stage Object grasp stage Object transport stage Object release stage Fig. 12.5 Diagram calculating the position of the end-effector in space S3 according to Fig. 12.3a -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0 -0.5 0 V er tic al -Y (m ) Horizontal -X (m) 1 2 3 4 5 the coordinates (x5; y5). In 1-position: \u2212 0.25m \u2264 x5 \u2264 \u22120.20m and \u2212 0.1m \u2264 y5 \u2264 0m; (12.7) In 2-position: \u2212 0.30m \u2264 x5 \u2264 \u22120.25m and \u2212 0.2m \u2264 y5 \u2264 \u22120.1m; (12.8) In 3-position: \u2212 0.35m \u2264 x5 \u2264 \u22120.30m and \u2212 0.3m \u2264 y5 \u2264 \u22120.2m; (12.9) In 4-position: \u2212 0.40m \u2264 x5 \u2264 \u22120.35m and \u2212 0.4m \u2264 y5 \u2264 \u22120.3m; (12.10) In 5-position: \u2212 0.45m \u2264 x5 \u2264 \u22120.40m and \u2212 0.5m \u2264 y5 \u2264 \u22120.4m; (12.11) With each COM coordinate of the manipulator (xcom; ycom) we calculate: \u23a7\u23a8 \u23a9 \u22120"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003964_elan.1140080709-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003964_elan.1140080709-Figure6-1.png",
        "caption": "Fig. 6. Typical cyclic voltammograms obtained with HMDE in 02-saturated aqueous solutions containing 0.1 M NaOH and 0.5 M KCI in thc presence of each surfactant (saturated): A) 3, I3) 4, C) 5, D) 6 , E) 7. In each case, the dotted cyclic voltammogram was obtained in the presence of IOpM SOD. Potential scan rate: 400 mV s-\u2019 .",
        "texts": [],
        "surrounding_texts": [
            "F. Matsumoto et a/. 650\nB\nFig. 4. A) Cyclic voltammograms obtained with HMDE in 02-saturated aqueous solutions containing 0.1 M NaOH and 0.5 KCI in the absence (1 j and the presence of both 1 ( I 7 mM) and 2) 10, (3) 20 and (4) 40 vol % I-propanol. Potential scan rate: 500mV s- . B) Cyclic voltammograms obtained with HMDE in 02-saturated aqueous solutions containing 0. I M NaOH and 0.5 M KC1 in the absence ( I ) and the presence of (2) 5, (3) 9 and (4) 17mM of 1. Potential scan rate: 500mVs-'.\n(17 mM) and different concentrations of I-propanol. The surfactant 1 is much more soluble in 1-propanol than in water. Increasing the concentration of 1 -propano1 at a constant concentration of 1 tended to shift the electrode reaction of O2 from the one-electron process of O2 to 0; (in the absence of 1-propanol) to the two-electron process of O2 to HOT. The cyclic voltammograms at various concentrations of 1 (Fig. 4 B) showed that the higher the concentration of 1 is, the smaller the redox response for the two electron process is and the larger that for the one-electron process is. These facts demonstrate that surfactant molecules of 1 adsorb on the HMDE, being in equilibrium between its surface and in the electrolyte solution and that as the amount of the adsorbed quinoline is decreased, the electrode reaction of O2 shifts from the one-electron O2 4 0, process to the two-electron O2 4 HOT process.\nFrom Figure 3 B, it i s obvious that the i,\"/i,\" value for the O,/O; couple decreases with decreasing u. This is presumably a result of the spontaneous disproportionation (or dismutation) or 0, to O2 and hydrogen peroxide (Eq. 3), although its effect is fairly small compared to those in neutral and acidic media [S, 32-37].\n202 + H20 + HO; + 0 2 + OH- ( 3 )\nThis presumption could be easily confirmed by adding a small amount of SOD enzyme into the same solution as that used in Figure 3 B, because SOD is an extremely efficient and selective catalyst for this disportionation [18, 19, :!4-27, 371. The cyclic voltammograms (Fig. 3 C) represent the typical result of such an experiment. In the presence of SOD (10pM), the cathodic peak current became larger compared to that without SOD and the anodic peak current could not actually be observed except with high potential scan rates. This result is in accord with expectations based on Equation 3.\nThe electrogeneration of 0; at the 1-adsorbed mercury electrode could be also confirmed by the reaction with ferric cytochrome c (Cyt.c(Fe3+)) in which 0, reduces Cyt.c(Fe3+) to its ferrous form [38-401\nCyt.c(Fe3+) + OT - Cyt.c(Fe2+) + O2 (4)\nIn the presence of Cyt.c(Fe3+), we could, as expected, observe the increase in ip\" for the one-electron reduction of O2 to 0 7 i.e., the catalytic current (see Fig. 5) . The reaction of Equation 4 has been widely used for in vitro 0; assay [41].\nThese facts confirm previous results [4, 6,20, 221 and strongly demonstrate that 0, can be electrogenerated by the oneelectron reduction of O2 at the 1-adsorbed mercury electrode in\naqueous media. In the following experiments to search for new surfactants for the electrogeneration of O;, we will examine by cyclic voltammetry,' the SOD-catalyzed disproportionation of OF which might be formed because of its unique specificity and high rate (it is almost diffusion-limited).\n3.2. New Surfactants\nAccording to the previous studies [I-71 and by considering the various reactivities of 0, such as basicity, nucleophility, free radical properties, redox properties, and so on [17, 18, 32, 42-45], recently we have found some surfactants for the 0, electrogeneration. The cyclic voltammetric results in their presence, which enabled us to reasonably diagnose the 0; formation, are shown in Figures 6 and 7. The data for 2 and 10, which have been reported previously by other researchers [ 1-3, 5-7, 24-27], are also given for comparison. All the surfactants shown here are those in the presence of which 0, could be actually electrogenerated at the HMDE, as can be seen from the following facts: (i) the redox response could be observed around -0.31 V and was obviously different from that (-0.14V) for the 02/HO, couple [3,20,32,46], (ii) the cathodic peak current ti,\") was about one-half that in their absence (compare with Fig. 3 A), and (iii) ip' was increased and no anodic peak current was",
            "actually observed in the presence of SOD. In addition, it is obvious that in the absence of SOD ii and i,\u201d/i,\u201d largely change with the surfactant, for example (see Table l) , ii increases in the order of 4, 5, 6, 7 < 3 < 1, 2 < 8 < 10 < 9 and i,\u201d/i,\u201d decreases in the following order: 1, 2 > 5, 6, 7 > 3, 4 > 8 > 10 > 9. This may primarily suggest the surfactant dependence of the electrogeneration of 0,. However, it is not easy to draw molecular factors of the surfactants desirable for 0,\nElectroanalysis 1996, 8. No. 7",
            "electrogeneration from these values of ii and i:/ii, because these both depend on the effectiveness of the surfactant matrix, which is built on the electrode surface, for the electrogeneration of 0; as well as the reversibility of the 02/0, electrode reaction itself at each surfactant-adsorbed electrode and the spontaneous disproportionation of 0;. Here, we will tentatively consider i,\"/i,\" as a rough measure of 'net effectiveness' for the electrogeneration of 07, which should result from all of these contributions to i:/ii. Based on this idea, we can see that the hydrophobic films built up by 1(2) and its methyl derivatives are more effective for the electrogeneration of 0, than those of 8 , 9 and 10, and that the effectiveness depends on the position of the methyl group introduced on the quinoline ring and decreases in the following order: 1, 2 > 5, 6, 7 > 3, 4, probably reflecting the differences in the molecular orientation and the amount of the surfactant adsorbed, the compactness and thickness of the adsorbed film, etc.\n3.3. Structure and Properties of the Surfactant Film Desirable for 0; Electrogeneration\nGierst et al. [31, 47-51] suggested that the molecular reorientation of several surfactants of the quinoline group occurred as a function of the electrode potential ( E ) and of the surfactant bulk concentration. Their results have been supported by the ellipsometric [30] and in situ FTIR spectroscopic [52, 531 studies of the adsorption of quinolines at the mercury electrode. In the present study, 0 < E < --0.6--0.8V (vs. Ag/ AgCl) which is expected to be more positive than the transition potential at which the 2-D phase transition from liquid-like to solid-like structure occurs [3 1, 47-5 11, and the surfactant is saturated. Thus, according to the criterion and data of Gierst et\nal. [31, 47-51], the adsorbed layers of 1, 2, 3 and 7 on the HMDE surface are considered to be liquid-like hydrophobic films formed by the surfactant molecules standing up. Similar liquid-like films may be also assumed for other surfactants examined here. Therefore, the formation of such a liquid-like hydrophobic film, which may result from various interactions (e.g., hydrophobic-hydrophilic interactions as well as the .rr-interactions between the surfactant molecules and/or between the surfactant and the mercury electrode, and electrode field-dipole interactions) and is considered to be a polymolecular loose layer of mobile structure and composed of small micellar aggregates [30, 311, must be a key point for 0; electrogeneration using the present electrode system in aqueous media.\nFigure 8 shows the cyclic voltammograms for the redox reaction of the O,/O, couple in 0.1 M tetra-n-butylammonium perchlorate media of the surfactants, which are found to be effective for the 0, electrogeneration as mentioned above. It is obvious from this figure that the O,/O; redox reaction is almost reversible and the electrogenerated 0; is stable, as in the case of the common aprotic solvents such as AN, DMSO and DMF [9-161. The surfactant film adsorbed on the HMDE may be thus considered to function as a 'quasi-liquid' of the surfactant liquid itself, as pointed out by Gierst et al. [31, 47-51]. Similar voltammograms were not obtained with 8 and 9, since they are solid at room temperature.\nWe could obtain no clear evidence to suggest 0; electrogeneration at the HMDE in the presence of the common anionic, cationic and neutral surfactants (e.g., sodium dodecylbenzene sulfate, hexadecyltrimethylammonium chloride and Triton X). In addition, no 0; electrogeneration was observed at the electrodes other than the mercury electrode (e.g., glassy carbon, Au, Pt and ITO) in the presence of 1 and its methyl and\nElectroanalysis 1996, 8, No. 7"
        ]
    },
    {
        "image_filename": "designv11_14_0002162_978-3-030-48977-9_1-Figure1.3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002162_978-3-030-48977-9_1-Figure1.3-1.png",
        "caption": "Fig. 1.3 Example of commonly used machine configurations. (a) Induction machine. (b) Synchronous machine. (c) Switched reluctance machine",
        "texts": [
            " This section aims to provide the reader with an overview of technology trends associated with key components of the electrical drive as shown in Fig. 1.1. Most importantly the objective in each of the ensuing subsections is to identify important developments and trends of key drive elements such as the machine, converter, and controller. The primary electro-mechanical energy converter of the drive is the electrical machine, which must be controlled in accordance with the industrial processes in which the unit is deployed. Modern (rotating) electrical drives typically use one of the three electrical machine types shown in Fig. 1.3. These machine types are referred to as the induction (asynchronous), PM synchronous, and switched reluctance machine and are shown in Fig. 1.3 consecutively from top to bottom and will be discussed extensively in this book. Both asynchronous and synchronous machine configurations depicted in Fig. 1.3 are shown with the typical three-phase winding, which is located in the stator slots of the machine. Note that other machine types, including the brushed DC machine, are also still in use. Of the three configurations shown in Fig. 1.3, the induction machine is most commonly found in industrial drives. This can be attributed to the inherent robustness of the machine itself and the presence of tried and proven drive components which form the basis of a reliable drive. Above all, the emergence of fast, low cost digital processors, and micro-controllers has been instrumental in achieving this market position, given that these controller units are able to accommodate well established control algorithms such as field-oriented control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000183_tim.2019.2949319-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000183_tim.2019.2949319-Figure9-1.png",
        "caption": "Fig. 9. Details of the beam aperture region of the bent DCW with optimized pillar locations and one pillar modified to avoid intersecting the electron beam. Inset: example of the beam hitting a pillar in the unmodified DCW.",
        "texts": [
            " Bending the DCW near the electron beam tunnels requires that pillars be locally modified to prevent them intersecting the beam. An optimization of the position and shape of the pillars in the bend regions was performed to maintain simultaneously the electrical behavior of the DCW, satisfy the fabrication constrains, and permit the electron beam to pass. It was found that by positioning the pillars correctly, only one pillar needed modification to maintain the electron beam flow. A portion of this pillar on the line of the beam was cut away (Fig. 9), so that it had an approximately triangular cross section. Numerical simulations confirmed that that this modified pillar geometry did not affect the electromagnetic behavior. Two DCWs with bent sections were fabricated in telluriumcopper alloy. The two structures differed only in the number of periods in the interaction region: one had 30 and the second 50 periods. Having SWSs with different lengths available permits experimental evaluation of the dispersion curve. Fig. 10 shows a photograph of the 50 period DCW before the lid was attached"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003709_s0043-1648(96)07486-8-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003709_s0043-1648(96)07486-8-Figure6-1.png",
        "caption": "Fig. 6. Tangential traction of rolling cylinders.",
        "texts": [
            " According to Kalker [13] the slicing method works when the contact area is slender in the rolling direction, which is the case in a spherical roller thrust bearing contact. Journal: WEA (Wear) Article: 7486 Each slice contact is treated as two cylinders rolling together.When two cylinders of similarmaterial roll together and transmit a tangential traction without sliding, there may exist a region of stick as well as slip (see Carter [14] and Poritsky and Schenectady [15]). In the case of complete slip the tangential traction per unit length is given by 22mp y q9s 1y (7)y 2pa a A second tangential traction, q0, acting over the slice is added to q9 (see Fig. 6): 2 2a9 2mp (yqd) q0sy 1y (8)y 2\u017e /a pa a9 The total traction q is then the sum of q9 and q0. The tangential load per unit length, t, at each slice can be expressed as (see Johnson [16]) 2 japG tsmp 1y 1q (9)\u00b5 \u2265 \u00a5 \u22024mp(1yn) where j is the creep ratio or the relative difference in rigid body velocities at each disk. Eq. (9) is valid if 0Fa9/aF1. If sliding occurs over the whole slice, the following expression for Coulomb friction is used: tsmp (10) Due to the curved contact surface in a spherical roller thrust bearing, the rollers will undergo sliding in the contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000998_tmag.2019.2902428-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000998_tmag.2019.2902428-Figure6-1.png",
        "caption": "Fig. 6. Distributions of eddy-current loss density. (a) 1-D FEM. (b) Method A. (c) Method B. (d) Method C. (e) Method D.",
        "texts": [],
        "surrounding_texts": [
            "We investigated the Cauer circuit modeling methods of dynamic hysteretic property in iron loss analyses of practical electric machines as the post-processing of the main magnetic field analysis. As a consequence, the computational accuracy of the Cauer circuit models is almost the same as the ordinary 1-D FEM including the ratio of eddy-current loss to hysteresis loss, although the computation time of Cauer circuit models is much lower. Therefore, the Cauer circuit models are one of the hopeful options to estimate the iron loss of electric machines within the acceptable computational cost."
        ]
    },
    {
        "image_filename": "designv11_14_0003484_9781119711230-Figure14.1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003484_9781119711230-Figure14.1-1.png",
        "caption": "Figure 14.1 Destination-sequenced distance vector routing protocol.",
        "texts": [
            " When the cable network is not available in areas like rescue then ad hoc network is the most feasible way to communicate. Routing paths in mobile ad hoc networks can potentially contain multiple hops, with every node in the mobile ad hoc networks acting as a router. Considering the mobility of the wireless host available in the ad hoc network, each of the nodes are equipped with the capability of the autonomous system without the use of any centralized administration. The distance vector routing protocol illustrated in Figure 14.1 is derived from the Bellman-Ford routing mechanism. It is a table-driven algorithm Semi-Automated Parking System Using DSDV and RFID 249 in which each node periodically broadcasts the routing updates. DSDV protocol requires each mobile node in the network to transmit its own routing table to its current neighbors. In all the nodes, each and every node in the table has entries for the destination node in the table and makes the count of the number of hops required for reaching the destination [2]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001066_4564709-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001066_4564709-Figure6-1.png",
        "caption": "Figure 6: Surface response generated from a quadratic model in the optimization of pH and adsorbent dosage for azinphos methyl, chlorpyrifos, malathion, and parathion.",
        "texts": [
            " This is also evident from the fact that the plot of predicted versus experimental values of OPPs correlation coefficient is close to y=x, showing that the prediction of experiment is quite satisfactory. 3.2.2. Response SurfaceMethodology. Response surfacemethodology (RSM) was developed by considering all the significant interactions in the CCD to optimize the critical factors and describe the nature of the response surface in the experiment. Three-dimensional surface plots were generated from the model fit in order to visually describe the interrelationship between the levels of factors and the recovery patterns of the OPPs (Figure 6). These plots were obtained for a given pair of factors at fixed and optimal values of other variables. The obtained curves of the plots indicate that there was interaction between the variables. From the model fit, it is observed that there is an increase in extraction recovery as the adsorbent dosage increases, which is due to the increased number of adsorption sites. The pHof a solution is also an important parameter affecting both the charge and stability in the extraction of OPPs during the adsorption process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure11.1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure11.1-1.png",
        "caption": "Fig. 11.1 Design scheme of fixed space-distributed combined aerometric receiver",
        "texts": [
            "2) where Va = \u2013V\u2014vector of true airspeed of motion of unmanned single-rotor helicopter relative to the surrounding air environment. An important aspect of technology for controlling air parameters on board unmanned single-rotor helicopter is the building of the sensor controlling system based on a single (integrated) fixed receiver of primary information. The original fixed aerometric receiver of informative parameters of the aerodynamic field of vortex column of rotor of single-rotor helicopter is proposed in works [6, 8] (Fig. 11.1). The basis of construction of receiver is multi-channel flow aerometric receiver 1 installed on fuselage of single-rotor helicopter. In flow channel between screen disks 2 and 3, the total pressure tubes 4 are installed which perceive pressures Pi with use of which according to developed algorithms the value of velocity V and the angle \u03c8 of the direction of the incoming air flow in azimuthal plane are determined [9]. The angle \u03b1 of incoming air flow in the vertical plane in the parking before starting the power plant of single-rotor helicopter determines according to pressures P\u03b1i, P\u03b1i\u22121 and Pst"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002375_nme.6490-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002375_nme.6490-Figure2-1.png",
        "caption": "Figure 2: Deposition process modeling",
        "texts": [
            " This article is protected by copyright. All rights reserved. A cc ep te d A rti cl e In this section, a thermal model of the FDM process is discussed. The FDM process modeling is divided into a deposition process and cooling process. We develop a model to emulate deposition of filament during material extrusion process. The deposition process is driven by a G-code which is a series of machine instructions for toolpaths. In a virtual space, the deposition process is modelled as successive addition of voxels as shown in Fig. 2. First, a given geometry is discretized into voxels which are rectangular cuboids. Then, the voxels are ordered along the printing path specified by the Gcode. This is done by using an in-house Matlab script. Given a list of voxels and their centers, we search along the toolpath over time to determine what voxels are added at any given time step. To facilitate the search, the voxel centers are put into a K-D tree (k = 3). With a current print head position as dictated by the G-code, we use a nearest neighbor search to find the nearest neighboring voxel center and add it to the physical domain if the path lies inside the voxel\u2019s space"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001993_s00170-020-06278-7-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001993_s00170-020-06278-7-Figure9-1.png",
        "caption": "Fig. 9 States relating to two-pass grinding. a State after the first grinding. b State before the second grinding",
        "texts": [
            " Thus, the flute grinding angle meets the design requirements for the second flute angle. Afs \u00bc \u03b1E\u2212\u03b1 J \u00fe \u0394\u03b1 \u00f012\u00de With this method, the center coordinates, OWEx OWEy\u00bd , are the coordinates of the end disk wheel. However, the wheel center point usually refers to the front disk wheel. The following coordinate transformation is therefore required: OWFx \u00bc OWEx \u00fe hw\u2219cos\u03b2 F OWFy \u00bc OWEy \u00f013\u00de For two-pass flute grinding, the grinding mark and edge width are determined by the value of the Z axis wheel center coordinate and the rotation angle of the tool. Figure 9 a shows the initial flute point, S, on the Y axis after the first grinding and the angle between the Y axis and point J, which is labeled Aff. However, the disk wheel that grinds the joint point for the second pass does not meet point J in this position. Figure 8 shows that the contact angle is \u03b1J. The tool therefore needs to rotate by an angle of ARO to ensure the accuracy of the second pass. The Z axis value of the end disk wheel center is ZJ when the wheel grinding point is J. The wheel center coordinates,OWFz, can then be described according to Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002765_s11071-021-06327-0-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002765_s11071-021-06327-0-Figure12-1.png",
        "caption": "Fig. 12 Sub-pitch meshing",
        "texts": [
            " The variation curve of the meshing tooth number is shown in Fig. 10. It demonstrates that the number changes periodically. 3.2 Sub-pitch meshing In sub-pitch meshing, PT\\PZ, and rp = 0.261 m, ra = 0.324 m. The difference between the two is recorded asD \u00bc PT PZj j. Figure 11 presents normal force curves with the same simulation condition as equal-pitch meshing. According to the occurrence time of the left surface contact force, it is known that the track pin is in contact with this surface when it is about to leave the mesh. As detailed in Fig. 12, traction is transmitted only by the tooth which is about to disengage from the engagement. There exist four regular intervals (D, 2D, 3D, 4D) between the remaining tooth surfaces and corresponding track pins on the sprocket pitch circle. When the tooth for transmitting traction force is withdrawn from meshing, the next tooth are not yet meshed. Then, a slippage of the pin relative to the contact surface occurs. It results in an impact force when the pin contacts with the left surface, as illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002137_j.ifacol.2021.04.127-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002137_j.ifacol.2021.04.127-Figure11-1.png",
        "caption": "Fig. 11. Prototype of case study 3D printer.",
        "texts": [
            " 10c, for example, all are mating interfaces in which 15 parts printed as a single part, while in Fig. 10a requires motion as the screw rail is replaced with the sliding function. It shows that functions of components are decomposed into feature-level functions. As in Fig 10a requisite to satisfy three functions: constrain DOF torsion of screw rail, rotational motion and constrains linear motion. These are all the vital details which are retrieved from the old design including the system boundaries to the new simplified design. Fig. 11 shows the prototype of the 3d printer, printed in our lab by using FDM process. Overall structural design is about modular design, in which simplification of design was attain through the accuracy and improved carrying capacity, free selection of different size drive modules, free choice of printing space and minimizing the risk of building a model incorrectly. Although, the above mention achievement was attained, there are still unsolved issues with function description makes it difficult to encode functional information in a uniform way"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003626_piae_proc_1922_017_035_02-Figure19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003626_piae_proc_1922_017_035_02-Figure19-1.png",
        "caption": "FIG. 19. FIG. 20.",
        "texts": [
            " 17 aic realised in oertain damped governors of the old Pamoins type and ir, an approximate way in the ,well-lubrioated springs of high-grade oars when newly assembled. There are one or two important variatio,ns on these main types. Conbider, for iiistanae, locomotive springs, which usually rim dry aid have very little nip. The friction b'etween the plates is of the solid type and varies with the load. The diagram is of the at The University of Auckland Library on June 5, 2016pau.sagepub.comDownloaded from 502 THE INS'I'I'I'U1'ION O Y A U1'O~IOHILE ENGINEERS form shown in Fig. 19, and this agrees with the esperimenk of Professor Dalby on Locomotive Springs.\" T h n , again, oonsider motor-car springs, which, for obvious masons of exprienoe, are given much nip and are as a rule slightly better lubricated. The loop in this case is usually parallel with a more rounded oorner in the Ceturns as in Fig. 20. This difference in oharacter between motor-car springs and loconiotive springs does not appear to have been notimd before. The esplanation will be given in more detail later"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003626_piae_proc_1922_017_035_02-Figure29-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003626_piae_proc_1922_017_035_02-Figure29-1.png",
        "caption": "FIG. 29.",
        "texts": [
            " F = 0 wherein the sign of F depends on the direction of motion; when & is positive the positive sign is used before F. Substitute .c = X T F/c2 and the differential equation above becomes the soIution of which is measure time from the instant when x = A , say the moment of extreme defleotion to the right, then Then i f , is the mass of the bar the equation of motion is na% + c 2 x = 0 X = A cos (ct/ n~ + 6) = .Z k F/c' Tho return swing then takes place about the point z = F / c 2 and the The motion is plotted on a t i e base in Fig. 29 and the true amplitudes next forward swing takes place about the point x = -P/cz. measured from the position of equilibrium if there were no friction are:= A - F/P = - ( A - YF/Pj . ~ f = (A - 5F/c?) s3 = - ('4 - iF/cZ) x4 = ( A - 9F/c'), etc. Thus bhe amplitudes decrease in arithmetical progression. Now using Mr. Raillie's definition of decrement as the ratio of each amplitude to the at The University of Auckland Library on June 5, 2016pau.sagepub.comDownloaded from 539 YRINCIPLES OF VE,flICLE SUSPENSION",
            " n 1 Z G = 0 Putting the matter more gcnerally, tlic amplitude is a linear function of the time, and if the time is measured backwards from the moment of coming to rest we have x = CI t ; the decrement is thus 11 = ~- f t\u2019 where b is the periodic or semi-periodic time. This becomes 8 = 1 + b / t a hyperbola agreeing very well with Mr. Baillie\u2019s esperimental curves. The outstanding fact is that even in his well greased spring, there is evidently solid friction far in excem of the fluid friation. It may be remarked that vibrations under solid friction may be projected from a rotating vector the end of which describes a form of spiral. (See Fig. 29. and Phil. Xag.. July, 1922, p. 284.) ROWELI,. at The University of Auckland Library on June 5, 2016pau.sagepub.comDownloaded from 530 7 HE IXSTITUTION OF AUTOMOHILE ENGINEERS. SPRING FRICTION EXPERIMENTS. Nr. Baillie in his paper on Springs mentions his difficulty, in connection with the zero, in obtaining deflections and decrements accurately, and thin difficulty is a general one in all damped elastic systems, e.g., galvanometers, torsion balances, etc. Because spring friction is of fundamental importance in suspensions and a8 it can be determind inad easily hy means of oscillat,ions, the following method of sunnountiny Mi-"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001689_j.optlastec.2020.106325-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001689_j.optlastec.2020.106325-Figure3-1.png",
        "caption": "Fig. 3. The schematic diagram of the self-building rolling fatigue wear tester [19]",
        "texts": [
            " The specific working parameters were involved (Cu target K\u03b1 ray; working voltage: 50 kV; current: 200 mA; scanning speed: 4 deg/min). The scanning ranges of Untreated, carburized matrix and different Units were all 20\u00b0- 90\u00b0. A Vickers microhardness tester (Buehler, 5104, USA) was used to measure microhardness. Different test samples were measured under different mechanical loads. Untreated and Mr Unit at 300gf, Mc unit and Cht at 500gf, Mcr unit, Msic unit and Mwc unit at 1000gf. Wear test was completed by self-assembled rolling contact fatigue wear tester, as shown in Fig. 3[19]. The overlapping load in the test was 8.3 kg. The speed of the motor was 700 rpm and the wear time was 30 h. The wear test conditions were established to simulate the operation of the steering axle ball joint. The quality loss of the worn samples was given by comparing the differences between the mass of the samples before and after the test. The above data was measured by a precision electronic balance with an accuracy of 0.0001g. Fig. 4 shows the structure at a distance of 0.1 mm from the surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001945_icra40945.2020.9196866-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001945_icra40945.2020.9196866-Figure1-1.png",
        "caption": "Fig. 1. Golem-KRANG.",
        "texts": [
            "00 \u00a92020 IEEE 3967 Authorized licensed use limited to: University of Canberra. Downloaded on October 03,2020 at 13:20:00 UTC from IEEE Xplore. Restrictions apply. underactuated dynamics or non-holonomic constraints. Current limitations of this framework stem from the implicit assumption that every DoF of the robot can be directly controlled [11]. However, this is hardly ever the case outside of the serial manipulator arms. RMPs and RMPflow have yet to be implemented on underactuated systems, like the robot used in this work, KRANG [12] shown in Fig. 1 (which has a WIP base with two serial manipulator arms attached on top, and one degree of underactuation). In this work, we propose a variant of RMPflow capable of executing RMPs on a class of WIP humanoids while maintaining dynamic stability. Our design is based on a hierarchical control and dynamics separation scheme. The main idea is to split the dynamics into two parts: the dynamics of purely actuated DoF and the residual dynamics that describes the relationship between actuated DoF and unactuated DoF",
            " Other remaining manipulation tasks are specified as RMPs and combined using the pullback operation of RMPflow into a single policy on the configuration space. This pullback policy is finally realized in the null space of the top-priority CoM trajectory tracking task. Overall, our scheme guarantees the dynamic stability of the locomotion by using the task prioritization idea in WBC, and uses RMPflow with the remaining DoF in an attempt to achieve multiple manipulation tasks without the algorithmic singularity due to successive projections in WBC. In simulations, we show the efficacy of our approach by designing motion policies for KRANG in Fig. 1 to solve combined locomotion and manipulation tasks. We review the essence of RMPs and RMPflow, as they will be used to define and combine the task-space policies in our framework. Further details can be found in [11], [13] Riemannian Motion Policies (RMPs) were first proposed by Ratliff et al. in [10] as a language to describe motion policies defined on general task spaces. Its main idea is to treat the task space as a manifold and model motion as geodesics (paths with the shortest distance). Using this correspondence, one can naturally think of motion policies as generators of these geodesics, and then controls the desired behavior through designing the manifold\u2019s Riemannian metric; e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000023_ipdps.2016.51-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000023_ipdps.2016.51-Figure6-1.png",
        "caption": "Figure 6. The encircled robots decide to start the run states, based only on the shape of the marked subchain within their viewing range. The large arrow heads indicate the moving direction of the runs. This is the notation, we will use for marking a runner. R(S) identifies the runner/robot which currently has the run state S. In (ii), the encircled robot is the endpoint of a horizontally and a vertically aligned subchain at the same time. Here, we must start two runs, moving in both directions along the chain.",
        "texts": [
            " For this, we first introduce a certain state of a robot, which we call the run state. We call a robot with active run state a runner. Robots can achieve this state in two different ways: 1) (start run state) If the local subchain within a robot\u2019s viewing range has a certain shape, then the robot decides on its own to generate the run state. We say that such a robot starts a run. Based on the shape of the local subchain, the run state gets a fixed moving direction along the chain. A robot can start and store up to two run states at the same time. Fig. 6 shows how the local shapes must look like. 2) (move run state) A runner R(S) can move the run state S to its chain neighbor r\u2032 in the moving direction of S. We say that the run state has moved from R(S) to r\u2032, while its moving direction initially set in 1) always remains unchanged. Afterwards, r\u2032 is identified by R(S). Once a run state has been started in 1), 2) is executed in each of the following rounds. This means that the run moves along the chain at constant speed and in the initially settled moving direction",
            " Note that depending on whether the distance between S1 and S2 was odd or even, it happens that during this passing process at one time both runs are located at the same robot. Then, this robot handles both runs separately according to their movement directions. C. Parallelizing runs: Pipelining The total process of gathering needs the work of many good pairs. For sake of a short running time, we let some of them work in parallel. Every constant number of L = 13 rounds, all robots simultaneously check if they can start new runs (cf. Fig. 6) and if so, they do so. We will show that this procedure ensures that even if multiple good pairs are nested into each other, different good pairs will enable different merges. This is what we call pipelining. Fig. 10 shows an example of this. Because S1, S2 are moving in the same direction, we call them sequent runs, while relative to their moving direction S1 is located in front of S2. The distance between them equals the number of edges on the subchain connecting both. As runs are moving with constant speed, the runs of the inner good pair will meet and as the result enable a merge and stop, first",
            " The main difference to the description of Section III is that we now let runs move along such quasi lines. Fig. 11 gives an example of a quasi line. Definition 1 (quasi line). We call a subchain a horizontal quasi line, if the following points hold: 1) At least its first and last three robots are horizontally aligned. 2) All its subchains of horizontally aligned robots contain at least three robots. 3) All its subchains of vertically aligned robots contain at most two robots. In a Mergeless Chain, at both ends a subchain of Fig. 6 in a matching rotation or reflection occurs. (If the chain is not mergeless, then the subchains outside the quasi line\u2019s endpoints may also have other shapes than these.) The definition of a vertical quasi line follows analogously. In Fig. 11, the fat subchain connecting and including the fat robots at its endpoints is called a quasi line. Having introduced quasi lines we now have to analyze, how this affects the merges, runs and pipelining we have explained in Section III. Merges remain exactly the same as in the basic description (Subsection III-A), i",
            ", they are only performed with black subchains, solely consisting of strictly horizontally aligned robots as shown in Fig. 2. So, we can continue with Mergeless Chains and the analog to Subsection III-B. The main approach remains the same as in Subsection III-B: i.e., if in Fig. 2 the black subchain is longer than the robots\u2019 viewing path length, we use reshapement hops of runners for shortening this black subchain until a merge becomes possible. We start new runs at the same subchains as in the basic description (cf. Fig. 6). The only difference is that now these subchains are connected by a quasi line. Because of this, as one can see in the example of Fig. 14, the runs now have to move several steps along the chain until arriving at the endpoints of the subchain, bordered in the figure, which needs to be shortened for performing the merge. For this movement, depending on the local shape of the subchain, we require additional run operations: Fig. 12.a) is the basic operation, while b) and c) are the new ones. a) The runner and at least the next 3 robots are located on a straight line. Here, the runner first performs a diagonal hop, then moves the run to the next robot. b) The runner and only the next 2 robots are located on a straight line. Then, for 3 times the runners just move the run to the next robot without any diagonal hops. Afterwards, it is located at the target corner c. c) This one is needed at most once for a new run, if started at the subchain of Fig. 6.(ii). New runs always start pairwise at both endpoints of a quasi line. Good pairs of runs are defined analogously to Subsection III-B by the relative position, concerning the quasi line, of the outer chain neighbors of the endpoints of the quasi line (cf. Fig. 13). We will show that also with quasi lines, good pairs always enable a merge. And as before in Subsection III-B, not all run pairs are good pairs. Fig. 14 shows an example of a good pair on a quasi line: The bordered subchain will be reshaped for performing a merge",
            " In order to connect two horizontal or two vertical quasi lines without enabling a merge, they must be connected by so-called stairways. Stairways can have arbitrary length and are subchains of alternating left and right turns. Fig. 17 shows an example. All differently shaped connecting subchains would allow merges. If a horizontal and a vertical quasi line are connected, then this can also be done without any stairway. The above construction exactly leads to the run starting subchains (i, ii) of Fig. 6. In Fig. 17, such subchains are bordered by dashed curves. So new runs are always started at the endpoints of quasi lines. For being able to close the chain, there must exist both, horizontal and vertical quasi lines. We start on the left of Fig. 18 at a robot s where a vertical and a horizontal quasi line are neighbors. For simplicity, we assume that the stairways connecting the quasi lines in the figure are of minimum length and symbolize quasi lines by dashed line segments. The fat robots correspond to the fat robots of Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure22.3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure22.3-1.png",
        "caption": "Fig. 22.3 Equivalent analytical model of an exoskeleton during the lifting load process",
        "texts": [
            " The hands also perceive the load, but to reduce it, flexible connections redistributing the load on the exoskeleton power frame can be used. The load lifting process scheme with the usage of exoskeleton is shown in Fig. 22.2. In this case, the exoskeleton can assist a person when lifting a load, in particular, compensating for a part of the hip and knee joints torques. Note that in this case, the exoskeleton can be considered as a three-link mechanism in which the force load is applied at the sling attachment point O4 (Fig. 22.3). The load force FL = MLg acts on the three-link system in such a way: at the point O4(ML\u2014mass of the cargo). The links are affected by given moments, which can be created by either passive (springs, brakes, etc.) or active (electric, pneumatic, and other types of drives) elements. The position of the mechanism links is determined by absolute angles \u03c6i . The most important issue during exoskeleton modeling is the kinematical and force parameters of the device movement determination that allows to achieve the required load movement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003720_s0039-9140(96)02040-1-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003720_s0039-9140(96)02040-1-Figure11-1.png",
        "caption": "Fig. 11. Capillary microcell for the batch injection analysis.",
        "texts": [
            " Although this effect was applied in macrocells, it can also be used in microcells to decrease the influence of the current which is not connected with the diffusion of the analyte to the working electrode surface. Hence it allows the detection limit to be decreased. In these microcells the working electrode can be in the usual state [5, 36, 37], but more often it is in the inverted state [6-10, 25, 94, 95]. The first microcell for voltammetric batch-injected analysis was proposed by Karolczak et al. [25]. This cell (Fig. 11) is distinguished by its exceptional simplicity. In Fig. 11, the working carbon paste electrode (1) in a Teflon sheath is put in a Teflon capillary (2). The sample (3) does not wet the Teflon and the carbon paste. The injection port (4) serves as an inlet for the sample (1-300 /~1) by a micropipette and for sample removal by an aspirator without washing. The residuals are negligible ( < 1%). The auxiliary electrode (5) is a platinum wire and (6) is the reference electrode in the form of a silver wire. A different approach to batch injection analyses was developed by Wang and co-workers [5-8], Amine et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001526_s00170-020-04987-7-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001526_s00170-020-04987-7-Figure1-1.png",
        "caption": "Fig. 1 a Different inherent strain calculation modes. b Triangulated surface representation of the part (STL) with the Z-axis corresponding to the build direction. c, d Voxel-based mesh with different voxel sizes",
        "texts": [
            " Additional features available in the software include the autogeneration of stress-based support structures, the detection of potential recoater blade crash events, the identification of high-strain areas, and the autogeneration of distortioncompensated geometries. While a summarized description of AP is presented here, more details can be found in the user\u2019s guide (ANSYS additive user guide, 2018 [26]). The modeling approach used in this software is based on the inherent strain\u2013based method in a layer-by-layer formulation. Three types of simulation (called strain modes) are available in the software (assumed uniform strain, scan pattern, and thermal strain), with three different approaches to calculate inherent strains. Figure 1 a presents the differences between each strain mode with a schematic representation of voxel inherent strain amplitudes using a color scale. In the assumed uniform strain mode, a strain value is uniformly applied to every voxel of the same layer, and to each layer of the part. In the scan pattern strainmode, an anisotropic strain value based on the actual scan orientation of the machine is uniformly applied to every voxel of the same layer. Finally, in the thermal strain mode, thermal simulation first predicts the strain at each point of the part, based on the actual scanning strategy of the machine",
            ", simulation accuracy vs solving time). On the geometry side, the part to simulate must be oriented along the desired printing direction (Z-axis) and converted into a stereolithography (STL) format. The meshing method is based on voxelization, where the size of a voxel (or an 8- node hexahedral element) should be at least one fourth (\u00bc) the minimum feature dimension and/or determined through a mesh sensitivity study. Therefore, it is assumed that multiple physical layers are embedded within one voxel layer. Figure 1 b\u2013d depicts the meshing workflow. Such a voxel-based approach is also used by other commercial LPBF simulation packages such as Simufact Additive\u00ae (MSC Software Company) and Netfabb Simulation\u00ae (Autodesk). To provide representative results, inputs about the scanning strategy of a given LPBF system (layer thickness, starting layer angle, layer rotation angle, etc.), the mechanical properties of the material to be processed (Young\u2019s modulus (E), Poisson\u2019s ratio (\u03bd), yield strength (YS), etc.), and the support structure features (minimum overhang angle and support yield strength ratio) are required"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003136_01423312211019654-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003136_01423312211019654-Figure1-1.png",
        "caption": "Figure 1. Ship model in the horizontal plane.",
        "texts": [
            " The rest of the paper is organized as follows: in the second section, the nonlinear ship motion model is introduced; in the third section, the receding horizon nonlinear identification approach for a time-varying nonlinear ship motion model is developed; in the fourth section, to validate the approach, a free-running experiment is designed, and the results are presented; and the fifth section provides concluding remarks for the entire paper. In this paper, we investigate the horizontal motion of a ship (Skjetne et al., 2004). The horizontal ship motion model contains two parts (Cui et al., 2010; Fossen and Perez, 2009), as shown in Figure 1. The left panel shows a physical drawing of the ship motion model, and the right panel shows the coordinate frame of ship motion. Based on the above coordinate system, the kinematics equation of the ship can be expressed as the dynamic equations: _h=R c\u00f0 \u00deV \u00f01\u00de M _n+C n\u00f0 \u00den+D n\u00f0 \u00den+ g h\u00f0 \u00de= t +w \u00f02\u00de In Equations (1) and (2),M ,C v\u00f0 \u00de, andD(v) are the system inertia matrix, matrix of the Coriolis and centripetal terms, and nonlinear damping matrix, respectively, h= \u00bdX , Y ,c T ,X and Y denote the position of the ship in geodetic coordinates, c is the heading angle, V = \u00bdu, v, r T and u, v and r denote the ship surging speed, sway speed, and yaw angular velocity, respectively, in the hull coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002097_j.addma.2020.101826-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002097_j.addma.2020.101826-Figure1-1.png",
        "caption": "Fig. 1. Comparison of the global and material coordinate systems which relates the applied load to the orientation of the layers which form the specimen.",
        "texts": [
            " (10) and rearranging gives the plastic multiplier which is expressed as, d\u03bb = \u2202f \u2202\u03c3 : Ce : d\u03b5 \u2202f \u2202\u03c3 : Ce : \u2202f \u2202\u03c3 + HT (11) where HT is the hardening modulus expressed by, HT = ca 2 3 \u2202f \u2202\u03c3 : \u2202f \u2202\u03c3 \u2212 ( \u2202f \u2202\u03c3 cX + \u2202R \u2202p )( 2 3 \u2202f \u2202\u03c3 : \u2202f \u2202\u03c3 )1 2 (12) Since the plastic multiplier has been defined, the plastic strain increment can be calculated as: d\u03b5p = \u2202f \u2202\u03c3 \u2202f \u2202\u03c3 :Ce:d\u03b5 \u2202f \u2202\u03c3 :Ce: \u2202f \u2202\u03c3 + HT (13) The layer orientation dependent material coordinate system can be related to the global coordinate system by a rotation matrix as demonstrated in Fig. 1, which is an approach also applied by [49,50]. The schematic demonstrates how, with different build orientations (relative to the build plate plane and the laser beam of the AM equipment), the material coordinate system is rotated about the global coordinate system. This approach can be used to extract information on the components of stress and strain. As illustrated in Fig. 1, the global coordinate system (\u03b5x,\u03b5y,\u03b5z,\u03b3xy,\u03b3xz, \u03b3yz) can be related to the material coordinate system (\u03b51, \u03b52, \u03b53, \u03b312, \u03b313, \u03b323) by considering a rotation about the x-axis, \u03b5\u2032 = q\u03b5qT (14) which can be expanded by considering the appropriate rotation matrix, \u23a1 \u23a3 \u03b51 \u03b312 \u03b313 \u03b312 \u03b52 \u03b323 \u03b313 \u03b323 \u03b53 \u23a4 \u23a6= \u23a1 \u23a3 1 0 0 0 cos\u03b8 sin\u03b8 0 \u2212 sin\u03b8 cos\u03b8 \u23a4 \u23a6 \u23a1 \u23a3 \u03b5x \u03b3xy \u03b3xz \u03b3xy \u03b5y \u03b3yz \u03b3xz \u03b3yz \u03b5z \u23a4 \u23a6 \u23a1 \u23a3 1 0 0 0 cos\u03b8 sin\u03b8 0 \u2212 sin\u03b8 cos\u03b8 \u23a4 \u23a6 T (15) This leads to transformed components of strain, \u03b51 = \u03b5x \u03b52 = \u03b5ycos2\u03b8+ \u03b5zsin2\u03b8+ \u03b3yzsin\u03b8cos\u03b8 \u03b53 = \u03b5zcos2\u03b8+ \u03b5ysin2\u03b8 \u2212 \u03b3yzsin\u03b8cos\u03b8 \u03b312 = \u03b3xycos\u03b8+ \u03b3xzsin\u03b8 \u03b313 = \u2212 \u03b3xysin\u03b8+ \u03b3xzcos\u03b8 \u03b323 = \u03b3yzcos2\u03b8+ \u03b5zsin2\u03b8 \u2212 \u03b5ysin2\u03b8 (16) where \u03b3xy = 2\u03b5xy, \u03b3yz = 2\u03b5yz, \u03b3xz = 2\u03b5xz, \u03b312 = 2\u03b512, \u03b323 = 2\u03b523, \u03b313 = 2\u03b513",
            " The parameter identification steps are summarised in the following sections. The components which define the materials anisotropy (F, G, H, L, M, and N) are derived by considering data gained from tensile tests of different build orientations. This derivation is based on the work by [49, 50]. Firstly, the parameters F, G and H are related to the tensile strengths in the orthotropic principal directions (T1,T2,T3) and the reference direction (T0) which is taken in this study to be the vertical build direction where \u03b8 = 0\u25e6 (orientation based on Fig. 1), G+H = T2 0 T2 1 , F +H = T2 0 T2 2 , F +G = T2 0 T2 3 (18) During the experimental analysis by Agius, et al. [51], the loading was always applied in tension (\u03b5z), which results in Eqs. (16) and (17) being simplified to the following which relates the applied stress and strain (considering the loading direction) to the components of stress and strain in the material coordinate system, \u03b52 = \u03b5zsin2\u03b8, \u03c32 = \u03c3zsin2\u03b8 \u03b53 = \u03b5zcos2\u03b8, \u03c33 = \u03c3zcos2\u03b8 \u03b323 = \u03b5zsin2\u03b8, \u03c323 = \u03c3z 2 sin2\u03b8 (19) Using Eq. (19) and the equivalent yield stress in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003922_28.806046-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003922_28.806046-Figure1-1.png",
        "caption": "Fig. 1. Induction motor per-phase equivalent circuit.",
        "texts": [
            " From (9), (12) and, therefore, (13) (14) These relationships allow the rotor -axis equation to be expanded in a similar way to (5) and expressed in complex form as (15) Similarly, the -axis equation may be written as (16) Rewriting (10), (15), and (16) in phasor form gives (17) (18) where and (19) The rotor current in (17) and (18) can be referred to the stator by balancing MMF\u2019s. Thus, (20) Substituting this expression into the phasor equations and rearranging gives (21) (22) Now, consider the induction motor equivalent circuit of Fig. 1. Summing voltages around each mesh loop gives (23) (24) Equating coefficients of and in (21) and (22) with those in (23) and (24) gives (25) (26) (27) (28) Relative skew between rotor and stator has the effect of modifying the stator/rotor mutual terms in (5) and (11) which are scaled by the skew factor. When this is carried through the derivation given above, the input impedance takes a form that can be linked directly to an equivalent circuit. Two alternative circuits may be used for this purpose, as discussed in [4]",
            " In the calculation of the equivalentcircuit parameters, the mean value of over a period of this modulation is required. This can be obtained by carrying out Fourier analysis on the waveform to extract the component which varies at rotor frequency. B. Validation of the Equivalent-Circuit Parameter Calculations When the multislice model is used to determine equivalentcircuit parameters, rotor skew is automatically included. There is, therefore, no need for the use of skew factors, and the equivalent circuit remains as shown in Fig. 1 with , , and all redefined as \u201ceffective,\u201d which means \u201callowing for the effect of skew.\u201d Conventional circuit theory leads to the model shown in Fig. 2, where , , and correspond to the unskewed values, and skew is accounted for via the skew factor . Clearly, when saturation is ignored, both models should be identical, and so, for example, calculated with the multislice model should equate to calculated using an unskewed rotor. To make this comparison, a preliminary study was carried out on an unsaturated motor at a full-load slip of 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001769_lra.2020.3006796-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001769_lra.2020.3006796-Figure3-1.png",
        "caption": "Figure 3. The configuration of the 5-DoF slave manipulator, which is named Teleop-Man. To simplify the problem, the controlled frame is chosen as the wrist of the Teleop-Man. In this case, the position is only determined by the first three joints and the orientation is only determined by the last two joints. The missing DoF is a pure rotational DoF.",
        "texts": [
            " The perpendicular curve \u03b3\u22a5 (\u03b6) is given by \u03b3\u22a5 (\u03b6) = Rs exp ( \u03b6RTs R\u22a5 ) \u03b6 \u2208 [0, 1] . (26) Corollary 1.2. \u03b3\u22a5 (\u03b6) gives the shortest path from Rs to the curve \u03b3Rm,\u03ba (\u03b8) in SO (3). ALGORITHM IN CASE STUDY To better demonstrate the proposed perpendicular curvebased Incomplete Orientation Mapping (IOM) algorithm, we take the asymmetric teleoperation system, where the master subsystem can provide 6-DoF pose sensing and the slave subsystem has only 5 DoFs, as an example. The configuration of the slave manipulator, which is named Teleop-Man, is shown in Fig. 3. To simplify the problem, the controlled frame of the slave robot is placed on the wrist of the Teleop-Man. The first three joints of the slave robot determine the position and the last two joints determine the orientation. Due to the deficiency in DoF, the slave robot loses the rotation ability around the direction \u03ba, which is perpendicular to the axes of the 4-th and 5-th joints. The robotic hand, which is installed on the end-effector of the Teleop-Man, is controlled by a hand exoskeleton. In this paper, the teleoperation of the robotic hand is not involved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001312_tie.2019.2952780-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001312_tie.2019.2952780-Figure5-1.png",
        "caption": "Fig. 5. Driving mechanism of the electromechanical actuator: (a) Driving mode for the clamping system, (b) Driving mode for the spindle system.",
        "texts": [
            " Two modes of the epicyclic gear are considered for the two different intended operations of the electromechanical actuator, which are explained as follows: 1) Clamping mode: During the clamping mode of the electromechanical actuator, the clutch is coupled with the stationary section as shown in Fig. 4(b) and the ring gear of the epicyclic gear is kept fixed using a spline structure as shown in left side stick diagram of Fig. 4(c), which is called as the planetary arrangement, highlighted in Table I. Furthermore, the spindle is connected with the stationary section and the drawbar, using the spline structure. The clamping mechanism between the spindle and the drawbar is highlighted in Fig. 5(a). Because of this coupling, the arrangement comprising of the clutch, the ring gear, the stationary section, the spindle and the drawbar are restricted from any rotational motion. When the drive motor rotates and the rotational motion is transmitted to the motor side shaft, shown in Fig. 5(a), which makes the sun gear of the epicyclic gear set shown in Fig. 4(c) to rotate at the same speed. This rotational motion is transmitted to the carrier of the epicyclic gear, which rotates at a speed less than that of the sun gear. This is because of the speed reduction using the gear ratio as shown in Table I during the planetary arrangement of the epicyclic gear. This rotation of the carrier is further transmitted to the drawbar, which cannot rotate because of the locking created by the spline mechanism with the spindle, however converts the rotation into linear motion by means of the screw-thread structure of the drawbar and finally generates the clamping force. To summarize, using the clamping mode, the rotary motion of the drive motor is converted into a linear motion of the drawbar by the help of epicyclic gear, which creates the clamping force required by the chuck during the machining. The locked and the power flow segments during the clamping mode are highlighted with different dashed lines in Fig. 5(a). Mathematically, the operation can be expressed as follows, The driving force for the whole electromechanical actuator is provided by the AC motor whose torque can be modelled by using the standard equations of an induction motor: Tm = J d\u03c9 dt +B\u03c9 + TL (1) where, J is the angular moment of intertia of the driving system containing the motor and load, B is the frictional coefficient, \u03c9 is the angular velocity and TL is the load torque. The motor torque is transmitted to the motor side plate and hence to the sun gear of the epicyclic gear using a pulley-belt mechanism",
            " During the spindle mode, from the right side stick diagram of Fig. 4(c) where all the parts can rotate, the ring and the sun gear are coupled to rotate together. Thus, when the motor rotates, it will rotate both the sun and the ring gear at the same speed. Moreover, the carrier in between the sun and the ring gear will also rotate with the same speed. This whole process will make the drawbar and the chuck, attached to the carrier to rotate at the same speed as that of the drive motor. This operation, thus called spindle mode is illustrated in Fig. 5(b). Similar to the clamping mode, the locked and the power flow segments for the spindle mode are highlighted with different dashed lines in Fig. 5(b). 0278-0046 (c) 2019 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. Averting failure of the whole electromechanical actuator structure by predicting and correcting potential failure scenarios at the design stage, before the actuator is built, is a key design strategy. Thus, during the clamping mode operation of the electromechanical actuator, which involves the 0278-0046 (c) 2019 IEEE",
            " Compared with the conventional hydraulic model, the only big difference for the proposed model is the exclusion of the hydraulic cylinder which reduces the spindle length (A) and power transmission unit length (B), whose values are presented in Table II. With the exclusion of hydraulic cylinder, there is no power transmission unit diameter (I). Instead the power is transmitted using the drive motor which is coupled to the pulley belt. Except that, pulley coupling portion (E) is longer than the hydraulic model as the clutch mechanism, shown in Fig. 5 is included there. The height of the spindle and the housing (F) is larger in the proposed model as it needs space for the planetary gear and spindle, which also makes the spindle height (G) bigger than the hydraulic model. Moreover, the dimensions are selected by considering the manufacturability of the electromechanical actuator as an independent module which can directly be installed on an existing lathe bed. During the simulation, the thrust of 50 kN is applied directly on the drawbar part, which is considered to be the weakest among the different parts of the electromechanical actuator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001916_j.compstruct.2020.112959-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001916_j.compstruct.2020.112959-Figure1-1.png",
        "caption": "Fig. 1. Windshield panel in skeleton design.",
        "texts": [
            " The high costs of the material and the manufacturing process limit their use in high\u2010volume car body structures. Thermoplastic composites offer easy processing for complex geometries, a good cost/performance ratio and their recyclability [2]. To be able to be economically viable, a tailored use of the composite material must be ensured. A design taking this into account is the so\u2010called \u201cskeleton design\u201d. This design was developed within the MAI\u2010Skelett research project and demonstrated on a windshield panel shown in Fig. 1. It combines the advantages of continuous fiber\u2010reinforced thermoplastic (CFRTP) composites with those of injection molding into a highly integrative design manufactured in three process steps. First the CFRTP profiles are manufactured using pultrusion. Those are thermoformed into the final geometry and subsequently overmolded using injection molding in order to join separate profiles and reach required mechanical performance. This symbiotic fusion of different materials is necessary in order to exploit the lightweight construction potentials [3]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001945_icra40945.2020.9196866-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001945_icra40945.2020.9196866-Figure3-1.png",
        "caption": "Fig. 3. A WIP humanoid with a 5-DoF arm mounted on a 2-DoF torso with 1-DoF locomotion which is underactuated. The blue circle denotes the CoM location of the robot and the red horizontal line at the end-effector indicates the orientation of the end-effector.",
        "texts": [
            " systems with number of actuators less than the dimension of the configuration space), including the WIP humanoids of interest here, In this work, we overcome this challenge by identifying a fully actuated subsystem based on the dynamics properties of WIP humanoids. We show that, under mild assumptions, these robots, despite being underactuated, can still be controlled by the RMPflow framework through this subsystem. This insight allows us to directly work with this recent advancement in policy fusion to design complex behaviors for WIP humanoids. For a general WIP humanoid with n joints and d-DoF locomotion shown in Fig. 3, its dynamics can be written as A [ x\u0308 q\u0308 ] +h = B\u0393 (4) where x\u2208Rd and q\u2208Rn are the coordinates of the center of mass (COM) and the joints, respectively, \u0393 = [\u03c41, . . . ,\u03c4n] > \u2208 Rn denotes the actuator torques, A \u2208 R(n+d)\u00d7(n+d) is the physical inertia matrix (which is symmetric positive definite), h \u2208 R(n+d)\u00d71 describes the combined Coriolis and gravity effects, and B\u2208R(n+d)\u00d7n is the actuation matrix. The system overall has n+ d DoF, but the number of control inputs is only n. Therefore the system is underactuated",
            " In practice, we can choose \u03930 as the minimal norm solution to the above problem, which can be computed in closed-form in terms of Moore-Penrose pseudo-inverse. Because of the parameterization in (11), the final policy will always achieve the desired CoM balancing and horizontal motion behavior, while the role of \u03930 is to realize other control tasks using the remaining DoF. V. IMPLEMENTATION AND RESULTS We verify the proposed control framework in Section IV in MATLAB simulation on a simplified4 2-D KRANG model (Fig. 3) that has a 2-DoF torso mounted on a rigid wheel and a 5-DoF serial manipulator arm. This system has one degree of locomotion, which is underactuated as described in Section III-A. The high-level controller generates the CoM angle trajectory using the MPC-DDP optimizer implemented in [16] for KRANG. The low-level controller is generated by RMPflow as discussed in Section II-C. Two task RMPs based on GDSs are implemented to maintain desired endeffector position and orientation. Details on the attractor implementation can be found in Appendix-A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002222_j.surfcoat.2020.125371-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002222_j.surfcoat.2020.125371-Figure11-1.png",
        "caption": "Fig. 11. Ultimate tensile strength and elongation of the different specimen.",
        "texts": [
            " We used EPMA to map the chemical composition of the reaction layers to accurately identify the phase of the bright white layer. An isolated W layer existed in the regular cellular W2C/TiC interface, and a continuous W layer existed in the irregular W2C/TiC mixed zone in the outermost region of the W2C layer, as shown in Figs. 9 and 10. The average thickness of the W layer in the laser melt injected coatings was approximately 200 nm, and that of the laser-induction hybrid melt injected coatings (preheating/post heating, preheating) was 300 nm. As shown in Fig. 11, ultimate tensile strength (UTS) of the LMI-ed coatings, the pre-laser-induction hybrid melt injected coatings and the pre/post-laser-induction hybrid melt injected coatings is 616\u00b1 43MPa, 691\u00b149MPa and 824\u00b1 39MPa, respectively. Among all the tensile specimens, the pre-laser-induction hybrid melt injected coating have the highest UTS, and UTS of the pre-laser-induction hybrid melt injected and pre/post-laser-induction hybrid melt injected coating is 34% and 12% higher than that of the laser melt injected coatings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000668_tcst.2016.2601286-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000668_tcst.2016.2601286-Figure1-1.png",
        "caption": "Fig. 1. CAD schematic of a range extender with rolling torque compensation [4].",
        "texts": [
            " As the generator shaft is rotating in the other direction as the crankshaft and the inertia torque will oppose its rotational speed change, the inertia torque of the generator will act against the inertia torque of the crankshaft. The inertia torques for internal combustion engine (ICE) and generator (Gen) are given by MInertia,ICE = JICE \u03c9\u0307ICE (1) MInertia,Gen = JGen \u03c9\u0307Gen = JGen \u03c9\u0307ICE i. (2) By choosing JICE = JGen i , the inertia torques are equal and as they are in the opposite direction they cancel each other out. A Computer Aided Design schematic of a cranktrain with generator and rolling torque compensation is shown in Fig. 1. 1063-6536 \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. The torque in the gear contact zone can be derived from both sides. From the ICE side, the torque equates to MGC,ICE = MICE \u2212 JICE\u03c9\u0307ICE. (3) For the generator side, the gear contact torque is MGC,Gen = MGen \u2212 JGen\u03c9\u0307Gen. (4) As the forces in the gear contact are equal, the torque relation is given by MGC,ICE = \u2212i MGC,Gen"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000671_med.2016.7535922-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000671_med.2016.7535922-Figure1-1.png",
        "caption": "Fig. 1: Quadrotor armed with a two-link robotic manipulator.",
        "texts": [
            " In order to achieve x-y positioning control and to compensate the torque \u03c41 due to the interconnection with the manipulator, we modify the aforementioned relations as U\u03041 = U1 (32) U\u03042 = U2 + \u03bdyc\u03c8 \u2212 \u03bdxs\u03c8 + T1,x (33) U\u03043 = U3 \u2212 \u03bdxc\u03c8 \u2212 \u03bdys\u03c8 + T1,y (34) U\u03044 = U4 + T1,z, (35) with the additional control parameters \u03bdx and \u03bdy defined as \u03bdx = kd,x ( x\u0307\u2212 x\u0307des ) + kp,x ( x\u2212 xdes ) (36) \u03bdy = kd,y ( y\u0307 \u2212 y\u0307des ) + kp,y ( y \u2212 ydes ) . (37) and the additional torques T1,x, T1,y , T1,z defined as\u23a1 \u23a3T1,x T1,y T1,z \u23a4 \u23a6 = \u03c41 \u23a1 \u23a31/la 0 0 0 1/la 0 0 0 1 \u23a4 \u23a6BRO \u23a1 \u23a300 1 \u23a4 \u23a6 (38) It is important to state that equation (38) is valid only when the first degree of freedom of the manipulator is rotational. However, a similar result may be obtained for the prismatic case as well. For the purposes of this work we decided to study the aerial manipulation system illustrated in Fig. 1. The system consists of a slightly modified \u201cPelican\u201d quadrotor, which is a UAV model implemented within the \u201cRotorS\u201d simulator, with a two-link robotic arm attached on it. The quadrotor mass is given as ms = 1 [kg], the arm length is considered as la = 0.21 [m] and the moments of inertia are defined as Ixx = Iyy = 0.01 [kg \u00b7 m2] and Izz = 0.02 [kg \u00b7 m2]. As for the robotic manipulator, we consider that the links are identical with masses m1 = m2 = 0.1 [kg] and lengths l1 = l2 = 0.1 [m], whereas the moments of inertia about their joints are given as I1 = I2 = 1 3m1l 2 1 [kg \u00b7m2]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002110_nap51477.2020.9309696-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002110_nap51477.2020.9309696-Figure4-1.png",
        "caption": "Fig. 4. Displacements in studied models.",
        "texts": [
            " Moreover, the distribution of residual stresses in the vertical printed prototypes is uneven: the maximum stress level of about 500 MPa is typical for the first deposited layer, and the minimum of 390 MPa for the last layer, that is attributed to less thermal impact on the layer. The verification of the calculations was carried out according to the magnitude of the displacements of points in forming body during the manufacturing of real samples with dimensions corresponding to the calculated model. The error was about 10% to 18% depending on the area of the trajectory of the layers. The amount of displacement in the manufactured samples does not exceed 0.28 mm (in the area of rounding of faces of prismatic samples), that is the maximum for the hollow prism sample (Fig. 4), which allows to ensure the appropriate level of accuracy of geometric dimensions of printed parts and to minimize the need for further machining. Using the TIG method and powder filler material for layer-by-layer synthesis results in poor formation of the bead of the first layer. It is related to the fact that the powder that is freely poured on the substrate is a dispersed system. Authorized licensed use limited to: East Carolina University. Downloaded on June 21,2021 at 00:15:29 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001790_j.precisioneng.2020.06.014-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001790_j.precisioneng.2020.06.014-Figure3-1.png",
        "caption": "Fig. 3. (a) Picture of the vibration-assisted drilling setup where the support, vibrating and drilling units are shown; (b) scheme of the vibrating unit composed by the sonotrode and the LPBF Ti6Al4V sample.",
        "texts": [
            " After the heat treatment, the microstructure evolved in a lamellar mixture of \u03b1 and \u03b2 phases inside the columnar prior \u03b2 grains, whereas the porosities got closed. As it is well known that the strong anisotropic characteristics of the AM parts cannot be eliminated even after the heat treatment, particular attention was paid to perform the experimental campaign on the same site of the sample. The vibration clamp used for the drilling experiments comprehends three different units: (i) the support unit, (ii) the vibrating unit, and (iii) the drilling unit, as shown in Fig. 3a. The first unit fixes the vibrating clamp system to the micro-milling machine table, which moves on the X\u2013Y plane. Moreover, the aluminium support has the function of connecting the ultrasonic transducer to the vibrating unit. The sonotrode and the Ti6Al4V sample form the vibrating unit: the former supports the latter through a mechanical joint (Fig. 3b). At last, the drilling unit is composed by the drill tool and the mechanical spindle. The whole system was connected to a 35 kHz ultrasonic generator (Herrmann Ultraschall, Ultrapack digital control 1000 M PK). The sonotrode design was carried out through a modal analysis using the Finite Elements (FE) ANSYS software (Fig. 4). Basic design of horn consists in the determination of the various key design parameters like resonance length, gain, nodal point, anti-nodal point, amplitude distribution and stress distribution along the length of horn profile"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000489_2016-01-1560-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000489_2016-01-1560-Figure1-1.png",
        "caption": "Figure 1. Ball-screw/nut interface showing the elastic deformations of a ball along its contact normal caused by contact load.",
        "texts": [
            " A simple case study is presented to illustrate the significant differences between the proposed model and Mei et al.\u2019s model [10], both benchmarked against an elaborate high-order Solidworks\u00ae FE model. Conclusions and future work, which aims to incorporate geometric errors into the proposed low-order model, are discussed. CITATION: Lin, B. and Okwudire, C., \"Low-Order Contact Load Distribution Model for Ball Nut Assemblies,\" SAE Int. J. Passeng. Cars - Mech. Syst. 9(2):2016, doi:10.4271/2016-01-1560. 535 Consider the ith ball of a BNA, sandwiched between the screw and nut as shown in Figure 1. Let \u03b4S(i) and \u03b4N(i) represent the deformation of the ball at the ball-to-screw interface and ball-to-nut contact points, respectively, caused by the contact force Pi. The deformations and contact force can be related by Hertzian Contact Theory as (1) where CS and CN are the Hertzian constants of the screw and nut, respectively, determined by the material properties and geometries the ball and screw/nut. The reaction force Pi, together with all the other reaction forces P1, P2,\u2026, PNball from the Nball loaded balls in the BNA, cause the screw and nut to deform elastically"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001001_sta.2019.8717197-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001001_sta.2019.8717197-Figure1-1.png",
        "caption": "Fig. 1. Ball Bearing representation",
        "texts": [
            " FEATURE EXTRACTION Bearings are essential components in the IM that ensure low friction while supporting high loads. Several types of bearings are distinguished, namely, the ball bearing, where the ball-ring contact is punctual, and the cylindrical roller bearing where the roller/ring contact is linear. Indeed, ball bearings are the most used in the industry because they offer the best price / performance ratio. Ball bearings are composed of an outer ring (OR), an inner ring, a ball and a cage. In Fig.1, VO, VC and VI represent the linear velocities of the OR, the ball center and the inner ring, respectively. Db is the diameter of the ball, Dc is the diameter of the cage, and \u03b8 is the contact angle of the bearing. The five basic frequencies are the shaft rotational frequency fr, fC the cage frequency, fPBE the characteristic frequency that reflects the frequency of the passage of a ball through an OR defect, fPBI the characteristic frequency that reflects the frequency of passage of a ball by a defect of an inner race and fball the frequency of the ball"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000204_j.autcon.2019.102996-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000204_j.autcon.2019.102996-Figure2-1.png",
        "caption": "Fig. 2. Parameters: (a) connection system; (b) coin angle; (c) twist angle. [21].",
        "texts": [
            " The application of the optimization on these joints provide results with less material; however, the geometry tends to be more complex, and the minimum bounding box of printed material remains practically the same. Connections can also be fabricated with molds and later cast with any appropriate material [6,18]. A disadvantage may be that a complex surface may require many different molds. There are many other types of nodes for gridshells [12], but one that stands out is the standardized hubs combined with flat-end tubes [1,19]. The hubs and flat-end tubes connection system (Fig. 2a) apply to a diversity of projects, such as roofs, fa\u00e7ades, bridges, and foundations [1,19]. Although the flattened ends of the tubes are positioned so that the strong axis resists the bending moments, the connection is prone to node rotation buckling. Additionally, the height of the hub depends on the coin angle of the connecting tubes as described below. Both characteristics are conveniently avoided by the new connection system proposed in Section 3.3.1. There are three types of parameters present in this system: element length, coin angle, and twist angle",
            " In doublelayer domes, for instance, the difference between the lengths of elements in different layers can create a curvature. Although lengths can be different, designers recommend that the size of the elements are not remarkably different from adjacent ones to avoid buckling of longer members. Additionally, constructors consider unpractical elements that are too long [20]. This parameter is relative to the mesh; it does not relate directly to the connection component itself. 2.1.1.2. Coin angle. The coin angle is the angle between the hub axis and the element axis (Fig. 2b), which determines how the flat edges of2 A box with minimum measurements that encloses the entire geometry. the elements are cut. Each element has two coin angles; they can be perpendicular to the element axis at both ends or cut at a certain angle. With this parameter, it is simple to add diagonal elements connecting multiple layers without having to change the node. 2.1.1.3. Twist angle. The flat ends of the elements can align to the hub axis, and, depending on hubs positions, the flat ends of the same element can have an angular difference about the element axis. This angle is identified as the twist angle parameter (Fig. 2c). 2.1.1.4. Nodes. The nodes, or hubs (Fig. 2a), are prismatic, standardized, with a central hole longitudinal to the hub axis and notches distribute radially around the extremity of the hub for the elements to connect. The central hole allows for a threaded rod to pass through and fix the flattened ends of the elements with washers. The number of notches in each hub may vary, as well as multi-layered system details. This connection system can solve a variety of surfaces. The producer's website [1] shows how flexible the system can be. We use the software Rhinoceros [22] in combination with the plugin Grasshopper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003709_s0043-1648(96)07486-8-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003709_s0043-1648(96)07486-8-Figure5-1.png",
        "caption": "Fig. 5. A contact surface made up of a series of thin slices, two of which are shown.",
        "texts": [
            " The load per unit length, p, as well as the effective contact length, l, is calculated several times in an iterativemanner, until the specified load is achieved. For each disk the contact semiwidth, a, is given by Palm- gren [11] for steel on steel 1/2 py3as3.35=10 (6)\u017e /8r where 8r is the curvature sum of the two bodies in contact. In the analysis of the tangential compliance, the contact area is divided into thin slices (one slice for each disk in contact) parallel to the rolling direction, neglecting any interaction between them(seeFig. 5). Slicing theorywas initiated by Haines and Ollerton [12] in 1963. According to Kalker [13] the slicing method works when the contact area is slender in the rolling direction, which is the case in a spherical roller thrust bearing contact. Journal: WEA (Wear) Article: 7486 Each slice contact is treated as two cylinders rolling together.When two cylinders of similarmaterial roll together and transmit a tangential traction without sliding, there may exist a region of stick as well as slip (see Carter [14] and Poritsky and Schenectady [15])"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002849_s13369-021-05659-8-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002849_s13369-021-05659-8-Figure2-1.png",
        "caption": "Fig. 2 Flux density distribution comparison of IM and LSPMSM designs",
        "texts": [
            " Electromagnetic analyses of the designed motors (IM and LSPMSM) were performed using the Maxwell software package. Experimental data from the IM was used to verify the FEA simulation results of the IM. This shows that the FEA simulation process is accurate and reliable and that the FEA simulation results for the baseline LSPMSM design can be considered reliable. The simulation results are based on operation at rated torque (26.7 Nm). The steady-state and starting performance of the LSPMSM are tested with different loading conditions using FEM. The flux density is given in Fig.\u00a02 and shows the maximum flux density value is approximately 1.87 ~ 2\u00a0T. This figure indicates that both motors safely operate under full constant load, with the LSPMSM having a slightly lower peak flux density than the IM. Figure\u00a03 represents the FE calculated torque response of both motors under constant full-load. The overshoot in the torque response curve of the LSPMSM is higher than that of the IM during start-up because of the magnet generated torque of the LSPMSM at the starting instant. The start-up of the IM is smoother and faster than the LSPMSM due 1 3 to the lack of PM braking torque (for the same frame size and weight)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001357_j.oceaneng.2019.106812-Figure19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001357_j.oceaneng.2019.106812-Figure19-1.png",
        "caption": "Fig. 19. Predicted cross section shape after rack and cylinder welding (deformed scale: 50).",
        "texts": [
            " Appling the computed temperature profile as thermal loading, mechanical analysis during welding was also carried out to compute the plastic strains, and predict the welding deformation and residual stress. Concentrating on the cross section shape of cylinder, Fig. 18(a) shows that the left edge has a negative welding deformation in horizontal direction, which means that the left edge expands outward with respect to origin. In addition, Fig. 18(b) shows that the upper edge also has a negative welding deformation in upright direction, which means that the upper edge shrinks inward with respect to origin. Dealing with the symmetrical characteristic, Fig. 19 shows the cross section shape of examined cylindrical leg structure after rack and cylinder welding, in which original and deformed shapes were marked with yellow and orange colors. It is also can be seen that same deformation trends comparing with experiment observation. Although the computed welding deformation has identical tendency comparing with experimental observation, it is still necessary to compare the magnitude of computed welding deformation and measured data. Computed Welding deformation of points on line 2 as indicated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001128_j.jfranklin.2019.08.018-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001128_j.jfranklin.2019.08.018-Figure1-1.png",
        "caption": "Fig. 1. Asymmetric actuator backlash characteristic.",
        "texts": [
            " In [54] , the backlash phenomenon was described by the inertia driven model, which was made up of two separated parts: the contact mode and the backlash mode. [28] investigated the active vibration isolation system, which included actuator backlash (a moving platform, linear actuators, a base platform, and external vibrations). In this paper, the backlash is a standard straight-line border backlash operator. The actuator backlash with input v b (t ) and output u ( t ) can be depicted by upward straight line (right) and d i u w a c u 2 w v w d ( ownward straight line (left), and they are connected with horizontal line segments, as shown n Fig. 1 , and the new model of the actuator backlash is represented as [8] (t ) = \u03d5( v b (t )) = \u23a7 \u23a8 \u23a9 m L [ v b (t ) \u2212 c L ] , m L [ v b (t ) \u2212 c L ] \u2264 u(t \u2212 1) m R [ v b (t ) \u2212 c R ] , m R [ v b (t ) \u2212 c R ] \u2265 u(t \u2212 1) u(t \u2212 1) , other cases (2) here m L and m R are the slopes of the two bold lines on the both sides of \u03d5-axis. c L and c R re the intersections of the v b -axis and the two bold lines. v b is the current actuator input. For simplicity, we introduce two variables s L = u(t \u2212 1) / m L + c L and s R = u(t \u2212 1) / m R + R "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002979_j.mechmat.2021.103951-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002979_j.mechmat.2021.103951-Figure1-1.png",
        "caption": "Fig. 1. The uniaxial loading-unloading test specimens a) As built SLM block, b) Final specimens according to ASTM E8M \u2212 04 standard.",
        "texts": [
            " While a predefined allowance for afterward machining was allocated during sample dimensions designing, all specimens were manufactured horizontally orientation. The specimens\u2019 dimensions and SLM processing parameters are listed in Table 1. The fabricated samples were placed in an oven at 650 \u25e6C for 2 h for stress-relieving. After stress relieving, the samples were removed from the base-plate, and the final shape of the specimens was obtained either by machining or wire EDM. The Loading-unloading test blocks, were turned into several test specimens according to ASTM E8M \u2212 04 standard (ASTM, 1991) as depicted in Fig. 1. It is worth to mention that, similar to reported observation in (Zhang et al., 2018a), the investigated cross-sections of foregoing blocks revealed no differences in the percentages of porosities occurred in the building direction. It means that the created samples should have identical microstructure. The final shape of the high-cycle fatigue test specimen are shown in Fig. 2. In order to quantify the manufacturing defects in produced specimens, optical microscopy was used to examine cross-sectioned characterization samples, see third row in Table 1, for any porosity occurring, For this purpose, the samples were cold mounted and burnished by K"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003751_s100510050161-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003751_s100510050161-Figure1-1.png",
        "caption": "Fig. 1. Schematic representation of the Hele-Shaw cell with lifting plates, showing the parameters used in the text.",
        "texts": [
            "-m Fluid surfaces and fluid-fluid interfaces Despite of the great effort devoted lately to improve the understanding of Saffman-Taylor (ST) instabilities [1\u20134], many related experimental facts still lack a sound explanation [3]. Here we are interested in the experimental observations on Hele-Shaw (HS) cells with lifting plates [5]. This variation with respect to the standard constant gap HS cell was suggested as a way to bring the ST problem closer to directional solidification [6,7]. In this experiment (see Fig. 1), instead of applying pressure to the less viscous fluid, the upper plate is lifted at the less viscous side (commonly air) at a fixed rate. It seems clear that the lifting of the upper plate will promote a pressure gradient analogous to the temperature gradient present in directional solidification. An interesting variation of this experiment is the Hele\u2013Shaw cell with a small gap gradient investigated by Zhao et al. [8] (see also Ref. [9]). The main experimental results obtained by Ben\u2013Jacob et al",
            " (2) Averaging over the the cell gap leads to the PoiseuilleDarcy equation for the mean velocity of the fluid, vj = \u2212 b2 12\u00b5j \u2207Pj , (3) where vj , \u00b5j , and Pj are the velocity, viscosity and pressure field of fluid j (j = 1, 2); and b is the gap of the cell. In order to obtain an equation for the pressure field we need to combine equation (3) with mass conservation. In the present case the latter deserves a careful consideration. Let the HS cell lie in the x\u2212y plane, the y\u2013axis being the direction of motion of the fluids, and the origin of coordinates be at the closed (fixed) end of the cell (see Fig. 1). The gap of the cell varies as b(y, t) = b0 + y tan(\u03c9t) , (4) where \u03c9 is the lifting angular speed (we shall hereafter call a = tan(\u03c9t)). As a consequence, the mass within a thin column of height b changes as \u03b4m/\u03b4t \u221d \u03b4b/\u03b4t (where the density of the fluid \u03c1 is assumed to be constant). Then, the equation which describes mass conservation reads \u2202b \u2202t = \u2212\u2207 \u00b7 (bvj) . (5) It should be here noted that mass (and density) conservation requires that, neither bubbles are formed nor drops of the displaced fluid are left behind during the lifting process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002564_wcsp49889.2020.9299760-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002564_wcsp49889.2020.9299760-Figure1-1.png",
        "caption": "Fig. 1. The sketch map of a UAVs group serving as removable BSs covering the ground users.",
        "texts": [
            " In order to keep the expansibility from single-agent to multiagent, agents only use their own local information in DGN. Each agent has its own DQN to obtain the policy while the input of the DQN has been through convolutional layers with relation kernel of multi-head attention [18], containing information of other agents\u2019 observations. A group of N UAVs are arranged to fly horizontally over a specialized target ground region. There are K ground users randomly scattered in the region, requiring to be covered by the communication signal coverage provided by the UAVs. As is shown in Figure 1, the UAV agents are aware of their own physical projection positions in the region, but know nothing about the positions of ground users in the beginning. As for ground users, the communication demands usually last for a period of time, so the time of each ground user being covered ought to be the main component of performance metrics. When UAVs work coherently as a team, it is economical and practical to utilize a decentralized structure since the fully centralized controlling consumes much communication resources"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002912_09544100211014754-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002912_09544100211014754-Figure3-1.png",
        "caption": "Figure 3. Structure of a planar free-flying two-link space manipulator postcapturing a noncooperative target.",
        "texts": [
            " Together with the definition of W \u00f0t\u00de, it follows that the system states qe, ve, bve, and b\u03b6 are all bounded during the time t2 \u00bd0,maxfT1,T2g . From the results of Steps 1 and 2, it can be obtained that the closed-loop system (22) is globally finite-time stable in the presence of lumped uncertainties. There exists a finite time T3 \u00bc maxfT1,T2g such that qe\u00f0t\u00de \u00bc 0n and ve\u00f0t\u00de \u00bc 0n, \"t \u2265 T3. This completes the proof. In this section, numerical simulations are performed on a planar free-flying two-link space manipulator postcapturing a noncooperative target (See Figure 3) to illustrate the proposed robust finite-time tracking control method. For convenience of the readers, the detailed dynamic equation of the two-link space manipulator is listed in the Appendix. The sampling time is set as Ts \u00bc 0:01 s. The physical parameters of the space manipulator are chosen as m0 \u00bc 60 kg, m1 \u00bc 6 kg, m2 \u00bc 5 kg, mt \u00bc 10 kg, I0 \u00bc 22:5 kg m2, I1 \u00bc 1:125 kg m2, I2 \u00bc 0:9375 kg m2, It \u00bc 2 kg m2, b0 \u00bc 0:75 m, a1 \u00bc 0:75 m, b1 \u00bc 0:75 m, a2 \u00bc 0:75 m, b2 \u00bc 0:75 m, a3 \u00bc 0:5 m, and qt \u00bc \u03c0=6 rad"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003687_a:1008389419585-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003687_a:1008389419585-Figure4-1.png",
        "caption": "Figure 4. Gear structure.",
        "texts": [
            " Two piezoresonators (accelerometers) with sensitivity 4.5 pc/g were placed and fastened on the base circles of gears in 180\u25e6 to measure the torsional vibration acceleration of gears. The acceleration signals of accelerometers were sent out and received by infrared rays devices. Then, the acceleration signals input computer by A/D transfer. The signals were processed and analyzed in time and frequency domains. There were two samples of the test gears, A and B, which had the same structure as shown in Figure 4. Each sample consists of two equal-sized gears. The specification of the samples of the test gears is listed in Table 1. The transmission error of the teat sample (A) shown in Figure 5 is decomposed into the random and the harmonic components, respectively, listed in Table 2. The harmonic components with frequencies 2\u03c0 and 70\u03c0 correspond to the eccentric error and the distance error between teeth, respectively. Figure 6 shows the autocorrelations of the second order Markov process and the random error components of teat sample (A)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001171_j.compstruct.2019.111423-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001171_j.compstruct.2019.111423-Figure3-1.png",
        "caption": "Fig. 3. Bending process of spacer monofilament. (a) When \u03b1=90\u00b0, tangent of point A is parallel to X axial. The surface of the fabric is tangent to point A. (b) The bending part of the monofilament becomes A\u2032B\u2032, AA\u2032 and BB\u2032 contact completely with the compression plate.",
        "texts": [
            " The compression process has resulted in a similar trend in the loaddisplacement curves for all the monofilaments, as shown in Fig. 2b. In the beginning stage of the compression (< 10% displacement), a larger load is required to break the stable state equilibrium with displacement in the monofilament, resulting in a high slope in the curve. A smaller load is needed for further compression with a lower slope as the monofilament has achieved the bending equilibrium. A relatively flat load-displacement curve is observed in the process of releasing compression, indicating the monofilament is still in bending equilibrium (Fig. 3). Finite element analysis (FEM) is an effective method in engineering analysis with high accuracy and great adaption to complicated shapes. The relationship of load-displacement can be acquired from the compression test with the purpose to estimate the load upon a specific displacement. Another issue is the initial shape of monofilament, in which the vertical monofilament may fracture under axial pressure ideally. An initial micro-bending is given to prevent compression fracture. The load-displacement graph is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure10-1.png",
        "caption": "Fig. 10. Scoop Tube Spray Formations",
        "texts": [
            " The cause of the loss of speed and the inconsistency as between one coupling and another was eventually traced to variations in the spray effect in the scoop tube chamber, and the resultant creeping of oil along the scoop tube housing at point X in Fig. 9, and thence down through the manifold drain ports into the reservoir. The leakage is prevented simply enough by fitting an anti-leak baffle as shown at point Y, but its occurrence was only discovered by cutting open the outer casing of a test coupling to study scoop tip effects under various running conditions. In the case of the coupling running clockwise, as shown in Fig. 10a, there was little spray and no need for the anti-leak baffle, since the \u201cbow wave\u201d of oil at the tip of the scoop tube is seen to be quickly merged with the annular ring of oil rotating with the casing. In the case of the opposite rotation shown in Fig. lob, however, the \u201cbow wave\u201d disturbance is seen to form a standing wave which collapses in the path of the oil entering the open scoop, and thus surprisingly reduces its capacity in terms of pressure and volume for a given degree of spray in the scoop tube chamber"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000486_1.4032993-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000486_1.4032993-Figure1-1.png",
        "caption": "Fig. 1 A disk with a sharp edge moving on a rough surface",
        "texts": [
            " Section 6 provides an explanation why a commercial toy could prolong its spinning time. Conclusions end the paper in Sec. 7. In this section, we first analyze the kinematics of the disk motion and employ the principle of virtual power (Jourdain) to derive the governing equations of disk dynamics, and then demonstrate how to incorporate rolling friction and Coulomb\u2019s friction into the dynamical equations to formulate an integrated model. 2.1 Kinematics of a Spinning Disk and Its Governing Equations. Figure 1 shows a steel disk with radius r, thickness 2h, and mass m in contact with an immobile table at point A. Let O and B be the disk\u2019s center of mass and the centroid of its bottom surface, respectively. A description of the disk\u2019s kinematics requires six degrees-of-freedom and two coordinate system frames: an inertial frame I \u00bc i; j;k\u00f0 \u00de attached to the table and a frame RO \u00bc O; e1; e2; e3\u00f0 \u00de fixed at the disk\u2019s center of mass O. The point O has the coordinates (x, y, z) in the frame I. The three unit vectors spanning the frame RO are chosen as follows: e3 is the axis of revolution of the disk, e2 is in the direction from A to B, and e1 \u00bc e2 e3 is thus parallel to the table"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002221_0278364919897134-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002221_0278364919897134-Figure11-1.png",
        "caption": "Fig. 11. Experimental setup for membrane inflation tests. The membrane was connected to a compressed air source and subjected to various pressures. An Arduino averaged data from a pressure sensor to provide pressure feedback. Two digital cameras, one above the membrane and another in front of the membrane, captured images of the membrane at each pressure step.",
        "texts": [
            " To accomplish this, the sample being tested, whether reinforced with fibers or not, was connected to a facility compressed air line with a simple, mechanical pressure regulator. As the applied pressure was increased by adjusting the regulator, a 20\u2013250 kPa absolute pressure sensor (MPXHZ6250AC6T1, Freescale Semiconductor) sampled by a 16-bit ADC (MAX1167 BEEE + , Maxim Integrated Products) controlled by an Arduino Mega 2560 provided an average value for the pressure over a 1 second interval. While the pressure was averaged, pictures were taken of the membrane from the front (to show the fiber pattern) and side (to calculate volume) using 8 MP digital cameras (Figure 11). The membrane was fixed to our setup for pressure testing using the clamping assembly shown in Figure 12. A quick connect tube fitting for 8 mm tube was threaded into a 0.25 in thick laser-cut (ULS PLS6MW, 50 W CO2 laser) delrin base. Blind holes were drilled into the base for dowel pins, which were placed at the end of each fiber to prevent the fiber from pulling the membrane out of the clamp. This also ensured that the fibers maintained the correct orientation during clamping and testing. A 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001790_j.precisioneng.2020.06.014-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001790_j.precisioneng.2020.06.014-Figure1-1.png",
        "caption": "Fig. 1. LPBF Ti\u20136Al\u20134V sample geometry (dimensions in millimetres).",
        "texts": [
            " The samples for the experimental drilling campaign were manufactured through Laser Powder Bed Fusion starting from commercially available Ti6Al4V grade 5 plasma-atomized powder (LPW Technology, UK). The material is a two-phase \u03b1 \u00fe \u03b2 titanium alloy, with aluminium and vanadium as alpha and beta stabilizers, respectively. The beta phase provides good mechanical properties, such as high strength and good ductility, while the alpha phase provides poor tendency to gases absorption. The chemical composition of the LPBF samples, evaluated using the Energy Dispersive Spectrometer (EDS) technology, is reported in Table 1. Ti6Al4V samples (Fig. 1) characterized by a working space (diameter of 22 mm and height of 10 mm) and a support pin for the coupling with the sonotrode (diameter of 5 mm and height of 8 mm) were produced on the MySint 100 AM machine (Sisma, Vicenza, Italy). The system is characterized by a cylindrical building volume of 780 cm3 (100 mm in diameter; height of 100 mm) and a laser unit having a maximum power of 200 W and a spot diameter of 30 \u03bcm. To maximize the material density, the samples were manufactured using the island scanning strategy with laser power of 105 W, scanning speed of 950 mm/s, hatch spacing and layer thickness of 80 \u03bcm and 20 \u03bcm, respectively. The build- M. Sorgato et al. Precision Engineering 66 (2020) 31\u201341 up direction was along the Z-axis (Fig. 1). After their manufacture, the samples were heat treated in a protective atmosphere of argon at 950 \ufffdC for 30 min, imposing a heating and cooling rate of 6 \ufffdC/min as recommended in the literature. The heat treatment is usually applied to LPBF parts to reduce porosities, stabilize microstructure as well as alleviate thermally induced residual stresses. The microstructure of the LPBF Ti6Al4V samples was observed by using a Leica DMRE optical microscope equipped with a high definition digital camera"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002137_j.ifacol.2021.04.127-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002137_j.ifacol.2021.04.127-Figure10-1.png",
        "caption": "Fig. 10. Function integration and structure simplification.",
        "texts": [
            " We replaced the screw rail as presently it cannot be printed using AM, segmentation printing of the housing floor plate as the length it too long to be printed by our printer, the external plate, simplifying and reducing the number of parts of the bearing the support. As a result, pointing out the functional features of the 3D printer, AMKR2 is applied first. Part consolidation vitally needs no relative motion, assembly access and single material between the component: for that reason, all parts are made as system of interests where functions are mapped to functional features as shown in Fig. 10. Arises to the fact that recognition of the system boundaries is primitive. As in Fig. 10c, for example, all are mating interfaces in which 15 parts printed as a single part, while in Fig. 10a requires motion as the screw rail is replaced with the sliding function. It shows that functions of components are decomposed into feature-level functions. As in Fig 10a requisite to satisfy three functions: constrain DOF torsion of screw rail, rotational motion and constrains linear motion. These are all the vital details which are retrieved from the old design including the system boundaries to the new simplified design. Fig. 11 shows the prototype of the 3d printer, printed in our lab by using FDM process. Overall structural design is about modular design, in which simplification of design was attain through the accuracy and improved carrying capacity, free selection of different size drive modules, free choice of printing space and minimizing the risk of building a model incorrectly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure6-1.png",
        "caption": "Fig. 6. A prototype of 6R Mechanism I.",
        "texts": [
            " 4 and 5 show Mechanism I at configurations A, B and C in circuit 1 and configurations D, E and F in circuit 2 respectively. It is observed that at two configurations in each circuit (see configurations A and C in circuit 1 and configurations D and F in circuit 2), the axes of R joints 1, 3, 4 and 6 of Mechanism I are coplanar and the axes of R joints 2 and 5 are perpendicular to the plane defined by the axes of R joints 1, 3, 4 and 6. The above results have been verified using several mechanism models built using 3D printing. Fig. 6 shows the CAD model and 3D-printed prototype of 6R Mechanism I (Fig. 4c). It is noted that joints 1 and 6 in this prototype are prevented from full-cycle rotation due to interference between links 2 and 4 as well as links 1 and 5. Let K1, K2 and K3 denote the intersections of joint axes of joints 1 and 6, joints 2 and 5, and joints 3 and 4. P1, P2 and P3 represent the plane defined by the axes of joints 1 and 6, joints 2 and 5, and joints 3 and 4 respectively (Fig. 7). From Ref. [36], we obtain that planes P1,P2 and P3 and plane K1K2K3 have a common point, K, at any configuration of the 6R mechanism during motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001079_1464419319862456-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001079_1464419319862456-Figure1-1.png",
        "caption": "Figure 1. Schematic map of the rotor bearing system with bolted flanged joints.",
        "texts": [
            " The solution of the nonlinear rotor system can be expressed as a periodic solution and other forms after bifurcation. To study the influence of bearing parameters on nonlinear vibration characteristics of the rotor system with bolted flanged joints, in this section, the IHB-AFT method is used to solve the rotor system, the periodic solution of the rotor system is tracked, and its stability is analyzed in a certain range of bearing parameters. A bearing rotor system with bolted flange joints is established as shown in Figure 1, two ball bearings are supported at the end of the shaft, where the two shafts of the rotor system are connected by a bolted flange joints. Assuming that the eccentricity of the disk of the rotor system is 0, the relevant parameters of the rotor system and ball bearing are shown in Tables 3 and 4. The finite element model of the rotor system with bolted flanged joints is established by the method described in the \u2018\u2018Finite element modeling\u2019\u2019 section. The equation of motion of the rotor system can be derived as follows M \u20acqr \u00fe \u00f0C ",
            " In the actual operation process, the clearance control of the rolling bearing can be realized by increasing the preload or counter-thrust of the high-pressure rotor system in the aero-engine, to eliminate or avoid the unstable motion interval caused by the nonlinear characteristics of the rolling bearing. Bearings are generally lubricated; contact stiffness is much lower than Hertzian contact stiffness. When the rotor system is in contact with medium and high loads, there is no bearing clearance at this moment, and Hertzian contact will occur.10 This section will discuss the influence of bearing clearance nonlinearity on the vibration characteristics of the rotor system. For the finite element model of the rotor system shown in Figure 1, the bearing clearance g is set to 5 mm, 10 mm, and 15 mm, respectively, the remaining parameters are shown in Tables 3 and 4. The amplitude\u2013frequency response curves of the rotor system with different bearing clearances are obtained when the angular velocity of the cage is 15 rad/s 4 4 80 rad/s. The results are shown in Figure 11. As shown in Figure 11, when the bearing clearance is 5 mm, the contact resonance occurred near the bearing cage angular velocity of 38.3 rad/s. With the increase of the bearing clearance, the performance of nonlinear characteristics of rolling bearings becomes more visible, for specific performances: The vibration displacement in horizontal and vertical directions increases obviously with the increase of the bearing clearance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003542_pime_proc_1948_158_045_02-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003542_pime_proc_1948_158_045_02-Figure4-1.png",
        "caption": "Fig. 4. Model Used in the Experiments",
        "texts": [
            " When a heavy mass charges into a light spring the combined system performs half an oscillation before separating. So, neglecting lateral friction at the treads in comparison with the inertia force, flange contact would last for 1/2N seconds, and the angle of yaw would then be (V/2N) x (h/2br) radians. On inserting in this expressioh likely figures for worn wheels at high speeds, an angle of the order of 0.01 radians is obtained, which is comparable with the maximum angle of yaw geometrically possible with standard flange and axlebox clcamnces. The Model. The model (Fig. 4) was one-tenth full-size, the wheel diameter being 4 inches. For the sake of compactness the r a i l s were circular rings, mounted on a drum whose spindle turned in fixed bearings. The wheels rested on the rails, and the wheel-and-axle unit had every degree of freedom except that of fore-and-aft translation. The arrangement was equivalent to keeping the axle stationary, and moving the track under it. The drum A was driven from a Ward-Leonard set, the tachometer W indicating the speed. The rail rings B had an outer diameter of 233 inches, a width of inch, and a depth of inch; they were attached to the drum by short spokes of spring steel, spaced, to scale, at sleeper pitch"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001447_s12555-018-0369-2-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001447_s12555-018-0369-2-Figure8-1.png",
        "caption": "Fig. 8. Process of NPGTI mode when \u03b7 = 90\u25e6.",
        "texts": [
            " migration caused by the Earth\u2019s rotation, the actual imag- ing point on the Earth surface in the next moment might be D2, namely the imaging point changes from D1 to D2, and this phenomenon should be compensated for through suitable attitude steering of the satellite. Thus, the drift angle is defined as the angle between the satellite projection scan velocity vp and the image motion velocity vi showed in Fig. 7, which can be determined from (24) as \u03b2 =arctan(vi1/vi2) =arctan ( |vD sin\u03b7 \u2212 vE sin(\u03b7 \u2212 i)| vp + vD cos\u03b7 \u2212 vE cos(\u03b7 \u2212 i) ) =arctan ( |\u03c9Rh cosbsin\u03b7 \u2212\u03d2sin(\u03b7 \u2212 i)| vp +\u03c9Rh cosbcos\u03b7 \u2212\u03d2cos(\u03b7 \u2212 i) ) , (25) where \u03d2 = \u03c9ERh cos\u03b4D and Rh = RE +h. In this paper we take the scenario in which \u03b7 is 90\u25e6 as an example shown in Fig. 8 [20]. Due to the earth rotation and the transport motion caused by the orbital motion , the actual sweep strip is a inclined one, not strictly perpendicular to the ground track, shown as the dotted line strips in Fig. 8. In order to realize the imaging of strips perpendicular to the direction of the ground track, the satellite needs carry out rolling control to compensate for the transport motion induced by orbital motion, and yaw adjustment to correct the migration caused by the earth\u2019s rotation. Based Fig. 8. Process of NPGTI mode when h = 90\u25e6. on aforementioned modeling analysis, the speed compensated for by rolling control is equal to the transport velocity caused by orbital motion, which can be expressed as vroll =vD = \u03c9 (RE +h)cosb. (26) The speed compensated for by yaw control is equal to the component of the earth\u2019s rotation speed vE in the direction of the orbital motion at point D, which amounts to vyaw =\u03c9E (RE +h)cos\u03b4D cos i. (27) Hence the steering laws of attitude angle for the satellite in NPGTI mode can be expressed as follow: 1) Roll angle According to the space geometry knowledge, the following geometric relationship is satisfied \u03b1 =arccos(cos\u03c6 cos\u03b8) , (28) where \u03b1 is the angle between the axis Zb of the bodyfixed frame and the axis Z of the orbital frame, and \u03b8 is the pitch angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001620_j.procs.2020.03.125-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001620_j.procs.2020.03.125-Figure6-1.png",
        "caption": "Fig. 6. Experiment setup",
        "texts": [],
        "surrounding_texts": [
            "In this section, we specify the equipment setup and the dataset of our system for the detection and classification of wrong strokes in table tennis."
        ]
    },
    {
        "image_filename": "designv11_14_0000067_metroi4.2019.8792876-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000067_metroi4.2019.8792876-Figure3-1.png",
        "caption": "Fig. 3. The optimized component for the antenna support.",
        "texts": [
            " MATERIALS AND METHODS The test case refers to a part, which is the support of the antenna of a satellite, in close collaboration with Thales Alenia Space as seen in Fig.1. This part started from an initial conventional design and was re-formed several times using topological optimization tools aiming to reduce the total weight of the component and maximize, in parallel, its rigidity. This process includes several steps like stress-analysis (external loads) and topological optimization (re-design) of it with respect to these specific loads. Both the original and the optimized components may be seen in Fig.2. and Fig.3, respectively. The material of this part is the Inconel 625 nickel-iron alloy which is frequently used in additive manufacturing process (PBF). The estimated weight of the initially designed part was 300 g while the final optimized design has an estimated weight of 270 g. The stress analysis and the topological optimization of the part are conducted using the Inspire SolidThinking software. Thus, using a certain procedure of iterative stress analyses and optimizations, the mechanical performance of the component is simulated and optimized"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002721_j.aej.2021.01.012-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002721_j.aej.2021.01.012-Figure6-1.png",
        "caption": "Fig. 6 Demonstrative figure of a machining worm model.",
        "texts": [
            " The position vector of the tooth flank surface of worm hob with the rectilinear axial profile is determined by r 1 0\u00f0 \u00de 1 \u00bc M 1 0 1 M12 M2 0 2 M21 x1 u\u00f0 \u00de y1 u\u00f0 \u00de z1 u\u00f0 \u00de 1 2 6664 3 7775 \u00f04\u00de After development, one gets the following form: r 1 0\u00f0 \u00de 1 u1; u\u00f0 \u00de \u00bc x1\u00f0u\u00de cos u1\u00f0 \u00de a sin u1\u00f0 \u00de \u00fe a cos u2\u00f0 \u00de sin u1\u00f0 \u00de\u00fe \u00fey1\u00f0u\u00de cos u2\u00f0 \u00de sin u1\u00f0 \u00de z1\u00f0u\u00de sin u2\u00f0 \u00de sin u1\u00f0 \u00de x1\u00f0u\u00de sin u1\u00f0 \u00de a cos u1\u00f0 \u00de \u00fe a cos u1\u00f0 \u00de cos u2\u00f0 \u00de\u00fe \u00fey1\u00f0u\u00de cos u2\u00f0 \u00de cos u1\u00f0 \u00de z1\u00f0u\u00de sin u2\u00f0 \u00de cos u1\u00f0 \u00de a sin u2\u00f0 \u00de \u00fe y1\u00f0u\u00de sin u2\u00f0 \u00de \u00fe z1\u00f0u\u00de cos u2\u00f0 \u00de 1 2 6666666666666666666664 3 7777777777777777777775 \u00f05\u00de In the case of the tool with arc axial profile, the position vector of the surface assumes the following form: r 1 0\u00f0 \u00de 1 \u00bc M 1 0 1 M12 M2 0 2 M21 x1 h\u00f0 \u00de y1 h\u00f0 \u00de z1 h\u00f0 \u00de 1 2 6664 3 7775 \u00f06\u00de After development, one gets the following equation: r 1 0\u00f0 \u00de 1 u1; u\u00f0 \u00de \u00bc x1\u00f0h\u00de cos u1\u00f0 \u00de a sin u1\u00f0 \u00de \u00fe a cos u2\u00f0 \u00de sin u1\u00f0 \u00de\u00fe \u00fey1\u00f0h\u00de cos u2\u00f0 \u00de sin u1\u00f0 \u00de z1\u00f0h\u00de sin u2\u00f0 \u00de sin u1\u00f0 \u00de x1\u00f0h\u00de sin u1\u00f0 \u00de a cos u1\u00f0 \u00de \u00fe a cos u1\u00f0 \u00de cos u2\u00f0 \u00de\u00fe \u00fey1\u00f0h\u00de cos u2\u00f0 \u00de cos u1\u00f0 \u00de z1\u00f0h\u00de sin u2\u00f0 \u00de cos u1\u00f0 \u00de a sin u2\u00f0 \u00de \u00fe y1\u00f0h\u00de sin u2\u00f0 \u00de \u00fe z1\u00f0h\u00de cos u2\u00f0 \u00de 1 2 6666666666666664 3 7777777777777775 \u00f07\u00de The parameter u1 in the above equations denotes the worm hob thread length. The parameter u1 varies from the initial value u1p to the final value u1k. Fig. 4 presents the surfaces of the globoid worm hob with rectilinear, concave, and convex axial tooth profiles. Three zones are distinguished on the tooth flank surface of the worm wheel (Fig. 5). Region II is the envelope to the family of contact lines of the globoid worm gear. Regions I and III are formed by the first cutting edge of the worm hob cutter (Fig. 6) [4]. One extreme cutting edge of the tool forms one side of the worm wheel tooth and the second edge forms the other flank. Region II of the worm wheel tooth flank surface is the result of linear contact of the worm and the generated worm wheel. The basic meshing condition is met, as shown below [4]: nxvx \u00fe nyvy \u00fe nzvz \u00bc 0 \u00f08\u00de where nx; ny; andnz are components of the normal vector to the surface, whereas vx; vy; andvz are components of the tangential vector. A relationship exists between rotations of the worm wheel blank and machining worm, which is represented by a homogeneous transformation matrix: M 0 2 0 1 0 \u00bc M 0 2 0 2 M21 M0 11 0 \u00f09\u00de Matrix M 0 2 0 1 0 in its development represents the following expression: M 0 2 0 1 0 \u00bc cos u 0 1 sin u 0 1 0 cos\u00f0u0 2\u00desin u 0 1 cos u 0 2 cos u 0 1 sin u 0 2 sin u 0 1 sin\u00f0u0 2\u00de 0 cos\u00f0u0 1\u00desin u 0 2 0 cos u 0 2 0 0 a cos u 0 2 a sin u 0 2 1 2 66664 3 77775 \u00f010\u00de xial profile: a) rectilinear, b) concave, and c) convex"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001798_s00419-020-01733-z-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001798_s00419-020-01733-z-Figure6-1.png",
        "caption": "Fig. 6 Schematic of designed test bench",
        "texts": [
            " 5 and 6, the schematic of the designed test bench, and also a section view of housing, a rotating shaft and bearings are shown. The two ends of the shaft are threaded to fix the positions of the inner race on the shaft by nuts and also to fix the position of the outer race on the housing by tightening the socket cap screws (please see Fig. 5). The electric motor and the rotating system are connected by a flexible coupling. The type of this coupling is \u201ccurved jaw flexible coupling\u201d which can absorb vibration, parallel, angular misalignments, and shaft end-play as shown in Fig. 6. The special grease KluberQuiet\u00ae BQ 72-72 which produces a considerably low-level noise is used for bearing lubrication. Before data collection, the environmental conditions are controlled and the necessary equipment is provided. The test bench is leveled to avoid applying axial loads, and the bearings and electromotor\u2019s axis are aligned with high-precision laser alignment equipment. All tests were performed at room temperature of 25\u00b1 2\u25e6 C that the room temperature was controlled by air conditioning system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000601_jnm.2157-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000601_jnm.2157-Figure1-1.png",
        "caption": "Figure 1. A 4P 60/48 WRindM for study.",
        "texts": [
            " In section 2, the case study is introduced and categorized as (i) turn functions and (ii) rotor and stator\u2019s slot opening and (iii) eccentricity fault. Section 3 deals with discrete machine modeling combined with modified winding function approach (MWFA) and in section 4, the simulation results are presented. Finally, section 5 concludes the paper. In this paper, in order to accurately compute the torque\u2019s ripple of a Three-Phase-Four-Pole WRindM, performance analysis considering the parameters listed in Table I is used (Figure 1). Floated star connection is considered for both rotor and stator windings. For simulation and modeling, a 4-Pole60/48-Slot WRindM has been considered as a study case. According to Figure 1, stator has 60 slots where five slots are assigned for each pole per phase. Rotor has 48 slots where four slots are assigned for each pole per phase. As shown in Figure 1, the reference of the stator circumambient (\u03b8s=0) and the reference of the rotor circumambient (\u03b8r=0) correspond to the center of winding (A) and the center of winding (a) respectively. Considering the number of Ns turns for each pole per stator\u2019s phase, the number of Ns/5 turns would exist in each slot. Also, considering the number of Nr turns for each pole per rotor\u2019s phase, the number of Nr/4 turns would exist in each rotor\u2019s slot. The machine\u2019s parameters and its structure have been shown in Figure 1. It should be noted that in order to use WFA, there are some parts including air gap and winding functions that should be modeled for dynamic modeling. Sinusoidal winding function is an ideal property of a machine that cannot be assumed for realistic and practical model because machines are not able to operate in their ideal status. As a matter of fact, limited number of slots is counted as one of the real factors that could lead to non-sinusoidal winding, turn and MMF functions [16\u201318]. Figure 2 shows the turn function of phase \u2018A\u2019 and \u2018a\u2019 Copyright \u00a9 2016 John Wiley & Sons, Ltd"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002020_s12206-020-1030-6-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002020_s12206-020-1030-6-Figure2-1.png",
        "caption": "Fig. 2. Bidirectional grasping mechanism: (a) grasping object from outside to inside; (b) grasping object from inside to outside.",
        "texts": [
            " The main difference between the presented four-bar finger mechanism and previous studies is that the presented four-bar finger mechanism aims to produce bidirectional grasping motion and increase the flexibility of object grasping. As depicted in Fig. 1(a), as the link AB (proximal phalange) is actuated and produce a rotation, the link BC (distal phalange) will rotate in the same direction. To produce bidirectional and symmetrical grasping motion, the length of link AB and CD should be designed equal. In this way, the gripper will be able to grasp the object from outside to inside (see Fig. 2(a)) as the most grippers act, like in Refs. [21-25]. In addition, due to the bidirectional symmetrical grasping ability, the gripper could also extend its finger from inside to outside and \u201cgrasp\u201d the inner surface of the object as depicted in Fig. 2(b). Due to the above characteristic of the finger mechanism, the flexibility of the gripper in object grasping could be enhanced to a certain extent. Then A gripper prototype composed of two finger mechanisms is developed to validate the effectiveness of the finger mechanism in bidirectional grasping. Due to the symmetrical movement characteristic of the finger mechanism, it will be beneficial to achieve modular design and flexible assembly. In addition, the number of the components of the gripper could be greatly reduced and result in lower manufacturing cost"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001532_0954406220908894-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001532_0954406220908894-Figure11-1.png",
        "caption": "Figure 11. Elastic deformations and stress contour of the runner block and the rail: (a) vertical compressive load; (b) vertical tensile load.",
        "texts": [
            " The magnitudes of the external loads are set as 1.5 kN. For the vertical flexibility, the proportion of elastic deformations is 31.95% when undertaking the vertical tensile load. While the proportion of elastic deformations drops to 15.02% when undertaking the vertical compressive load. For the lateral flexibility, this proportion is 31.67% with the vertical tensile load and 15.02% with the vertical compressive load. The elastic deformations of the runner block and the rail cause the stiffness characteristic differences discussed above. Figure 11 shows the elastic deformations and stress contour of the linear rolling guideway under vertical compressive load and vertical tensile load. The magnitude of the vertical load is set as 2 kN. With the vertical tensile load applied, the upper raceway grooves (raceway number 1 and 2) undertake larger loads. The structure characteristics of the runner block and rail determine that the elastic deformations are larger when undertaking the vertical tensile load. According equations (1) to (6), the elastic deformations have an effect on the contact conditions of the contact areas between rolling balls and raceway grooves"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001151_tdmr.2019.2937988-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001151_tdmr.2019.2937988-Figure2-1.png",
        "caption": "Figure 2. The 8-inch wafer with a polished surface on the backside, from which the test specimens with a size of 10 mm\u00d710 mm were cut with various thicknesses (t=105 \u03bcm, 82 \u03bcm, 57 \u03bcm, and 42 \u03bcm).",
        "texts": [
            " Hence, with the increase of this nonlinearity, the feasibility of the linear beam solutions for the bending test would not hold. Furthermore, to the best of our knowledge, there are no existing threepoint bending tests for dealing with this issue from testing standards such as JEDEC, ASTM or ISO. Therefore, the purpose of this study is to investigate the geometric nonlinear effect on the three-point bending test, and then to provide easy-to-use solutions for determining the bending strength of thin silicon dies. (a) Test Specimens The 8-inch wafer on a dicing tape is shown in Fig. 2 with polished surfaces on the back side. For the back-side surface preparation of this wafer, wafer back-side grinding was carried out by using 400 grit abrasive for a coarse 1530-4388 (c) 2019 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. grind first from the wafer thickness of 750 \u03bcm down to near the target thickness, then by using 6000 grit abrasive in a fine grind for last 40 \u03bcm thickness, and finally by polishing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003630_047134608x.w1111-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003630_047134608x.w1111-Figure1-1.png",
        "caption": "Fig. 1. Control effectors on a modern aircraft.",
        "texts": [
            " Relative wind refers to the instantaneous direction of the air mass that an observer moving with the aircraft would measure, that is, the direction of the air mass relative to the aircraft. Thus, in a quiescent air mass, one that is not moving with respect to the earth\u2019s surface, the direction of the relative wind is simply equal and opposite to the direction of the velocity vector of the aircraft center of gravity. In order to change the aircraft\u2019s orientation, the aerodynamic control effectors just mentioned are employed, and the magnitude and possibly the direction of the thrust force created by the propulsive system are varied. Figure 1 is a representation of the control effectors on a modern aircraft, here chosen as a tailless fighter. The control effectors in Fig. 1 are essentially devices to create aerodynamic moments about the aircraft center of gravity. The effect of these moments is to create rotational acceleration about the center of gravity 1 and hence change the orientation of the aircraft with respect to the relative wind. This, in turn, alters the total aerodynamic forces and moments acting on the aircraft. The changes in aerodynamic forces and moments created by the change in orientation are typically much larger than those created by the effectors themselves, and it is the former forces and moments that change the aircraft\u2019s trajectory and speed. As Fig. 1 indicates, tradition has led to the description of the change in orientation of an aircraft as pitching, rolling, and yawing motions, each described by a rotation with respect to an axis system fixed in the vehicle with origin at the center of gravity. The principal aim of aircraft control is to modulate the aerodynamic and propulsive forces so as to produce the desired motion of the vehicle. In the early days of flight, this was accomplished solely by the pilot, who determined the desired trajectory and speed and, through training, moved the control surfaces by means of hand and foot controls (referred to as cockpit inceptors) so as to produce the desired results",
            " Indeed, in most modern high-performance aircraft, the pilot is no longer directly connected to the control effectors by means of cables, pulleys, etc. Pilot commands are sent directly to the flight control computer, and the resulting system is referred to as fly-by-wire. The design and implementation of an aircraft flight control system must begin with an appropriate mathematical representation or model of the dynamics of the aircraft, that is, developing the equations that describe how an aircraft responds to control effector actuation and to atmospheric disturbances such as turbulence. Body-Fixed Axis System. Referring to Fig. 1, one sees an xyz axis system consisting of three mutually perpendicular (orthogonal) axes with origin at the aircraft\u2019s center of gravity. These so-called body-fixed axes are important in the development of the mathematical model of the aircraft. As the name implies, the bodyfixed axes are fixed in the aircraft body or airframe. The axis system has its origin at the aircraft center of mass, with the xz axes lying in the aircraft\u2019s plane of symmetry. In the equations to be presented, stability axes will used, wherein the x axis will be aligned with the relative wind when the aircraft is in equilibrium or \u201ctrimmed\u201d flight. Earth-Fixed Axis System. In addition to the body-fixed axes just defined, the description of aircraft motion and, in particular, its motion relative to the earth requires the introduction of an earth-fixed axis system, x\u2032y\u2032z\u2032, also shown in Fig. 1. Traditionally, this axis system has its z\u2032 axis directed toward the center of the earth, with the position of its origin and x\u2032 axis direction up to the analyst\u2019s discretion. For example, the origin may be assumed to lie at a particular geographical location on or above the earth\u2019s surface with the x\u2032 axis pointed in the direction of magnetic north. Equations of Motion. Equations Expressed in Earth-Fixed Coordinates. The equations that follow are similar to those to be found in any standard text on the subject, for example, Etkin (1), Nelson (2), or Schmidt (3)",
            " Accomplishing this requires the introduction and definition of three angles referred to as the Euler angles. These angles can be used to uniquely prescribe the orientation of the body-fixed axes relative to the earth-fixed axes through three ordered rotations as shown in Fig. 3. Note that the order of the rotations , , is essential to the definition of the Euler angles. Equations Expressed in Body-Fixed Coordinates. The vectors F, vc, G, and h can be conveniently expressed as components in the body-fixed xyz axes of Fig. 1 as follows: where i,j,k represent unit vectors parallel to the x, y, and z body-fixed axes and where X, Y, and Z represent the components of the aerodynamic and propulsive forces alone, with the contribution of the gravitational forces now included in such terms as \u2212mg sin . The scalar quantities L, M, and N are usually referred to as the rolling moment, pitching moment, and yawing moment, respectively. The angular velocity of the aircraft as evaluated in the earth-fixed axis system must also be introduced and defined",
            " The element in the ith row and jth column of G(s) is the transfer function between the ith output variable and the jth input, and the element in the ith row and jth column of H(s) is the transfer function between the ith output variable and the jth turbulence input. These transfer-function elements describe the relationship between inputs (or disturbances) and outputs. For example, given the state and output descriptions of Eq. (22), we have where \u03b8 is the aircraft pitch attitude (an Euler angle) and \u03b41 is a particular control input, perhaps the elevon input in the aircraft of Fig. 1. The roots of the numerator and denominator polynomials on the right-hand sides of Eq. (13) are referred to as the zeros and poles of the transfer function, respectively (9). In the absence of other inputs, the time response \u03b8(t) to the control actuation \u03b41(t) can then be obtained as where L \u2212 1 represents the inverse Laplace transform. Stability (i.e., bounded responses to bounded inputs) requires that the poles of the transfer function g11(s) lie in the left half of the complex s plane. There is a fundamental, or natural, mode of motion associated with each real pole and each pair of complex conjugate poles of the g11(s)",
            " Outer guidance loops with lower bandwidths are then closed about the inner loops. The sensor requirements in terms of signal-to-noise ratios for the outer guidance loops are less demanding than those for the inner control loops. With a control system like that just outlined in operation, an aircraft can be forced to follow a desired trajectory in three-dimensional space at some desired speed. In terms of mission effectiveness or economy some trajectories can be considered more desirable than others. For example, consider the fighter aircraft of Fig. 1 in a combat situation. To avoid an adversary the pilot must change heading (Euler angle ) by 180\u25e6 as quickly as possible. The question is: What trajectory and speed profile will allow this heading change to occur in minimum time? Next consider a large passenger aircraft such as that shown in Fig. 4. To minimize ticket costs it is necessary to minimize fuel consumption. The question now is: What trajectory and speed profile from departure point A to destination point B will result in minimum fuel consumption",
            " The design of the previous section was successful, given the single flight condition, the limited number of control inputs considered, the relatively simple design specifications, the underlying assumption of linearity, and the inherent stability of the basic aircraft. Even with these assumptions, however, the ability to move from a set of performance specifications to a final control law is not straightforward and certainly not algorithmic. Success is highly dependent upon the skill and experience of the analyst. Some modern flight control problems rarely allow the simplifications just enumerated. Consider the aircraft shown in Fig. 1. The absence of a vertical tail, dictated by a desire to minimize radar signature (i.e., to provide a stealthy design), means that at least one of the lateral directional modes of the vehicle will be unstable. To maximize maneuverability it is also likely that longitudinal modes will also exhibit instability. In addition, a wide variety of flight conditions will be contained within the aircraft\u2019s operational envelope. The use of the linear models in Eqs. (19) and (20) may not be justified in some flight regimes such as those in large-disturbance maneuvers",
            " That is, the eigenvectors determine how each state variable contributes to the vehicle response for each possible mode of motion. Like eigenvalues, open-loop and closed-loop eigenvectors can be defined. As a control-system design technique, eigenstructure assignment allows the creation of controllers that meet mode-based performance specifications. Examples of mode-based specifications are minimum damping, minimum settling time, and decoupled responses. The last refers to allowing specified system inputs to affect only certain system outputs. For example, in the case of the fighter aircraft of Fig. 1, it may be desirable to have the elevons affect only pitching motion, while the spoilers affect only vertical velocity. One of the attractive features of eigenstructure assignment for the control of piloted aircraft is the manner in which design criteria for handling qualities can be directly incorporated into the mode-based controller specifications. In addition, the order of the compensator is at the discretion of the designer. Possible disadvantages of this approach include the difficulty in determining how to tune an eigenstructure assignment design (improve its performance) if its performance in simulation or flight is unsatisfactory",
            " In other adaptive algorithms, the process of adaptation often involves system identification, wherein models of the vehicle are derived online, in real time. Aircraft applications of adaptive control are challenging in that the adaptation process as a whole must occur very quickly. For example, one of the most challenging current applications of adaptive control lies in the area of reconfigurable flight control. Here, the controller is required to adapt to damage to the aircraft or to subsystem failures. In the case of combat aircraft such as that of Fig. 1, wherein vehicle stability is dependent upon active feedback control, damage or subsystem failures can easily cause the vehicle to become uncontrollable by the pilot in a matter of seconds or less. As an example, Bodson and Groszkiewicz (22) offer a treatment of this problem using model-reference approaches. The attractiveness of adaptive control is its promise of providing control laws for aircraft applications in which significant and unpredictable changes in vehicle characteristics may occur"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002764_0142331221994393-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002764_0142331221994393-Figure10-1.png",
        "caption": "Figure 10. Car-like robot model.",
        "texts": [
            " However, this step will reduce the peak and average velocities of the car-like robot, thereby failing to satisfy the shortest time optimal solution. m is adopted to adjust the influence of the superimposed APF on the velocity in different scenarios. This key parameter determines the final velocity of the car-like robot. During the experiment, the reasonable velocity value is determined by continuously adjusting m. r0 is mainly combined with the repulsive potential field and dynamic potential field to determine the range of obstacles. A suitable r0 is determined through continuous adjustment to ensure stable robot movement. Figure 10 displays the car-like robot used in the practical test. This robot uses an LS01G lidar and an industrial computer configured with i5-6360U central processing unit and 4 GB random-access memory. The system is based on Ubuntu16.04+ROS Kinetic. The movement process of this car-like robot is shown in four parts (Figure 11). In part I of the actual test process, the car-like robot plans a smooth path and moves forward away from static obstacles (Figures 11(a) and 11(b)). In part II, a dynamic obstacle suddenly appears, and the front wheel of the car-like robot turns right to the limit position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002009_s42417-020-00259-6-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002009_s42417-020-00259-6-Figure5-1.png",
        "caption": "Fig. 5 Schematic diagram of discretized dynamic load and ring gear loading",
        "texts": [
            "65\u00a0Hz) and its frequency doubling component energy. As the frequency increases, the energy gradually decreases. Finite element analysis of\u00a0dynamic stress of\u00a0ring gear In an ideal state, during the meshing process of the planet and the ring gear, each participating tooth pair will always be in line contact. Therefore, to effectively simulate the planetring gear meshing process, the contact surface of the ring gear is divided into n contact lines along the direction of the tooth profileas shown in Fig.\u00a05b (Fdi is the impact load when the ith contact line participates in meshing), which will effectively reduce the mesh refinement degree of the contact surface of the ring gear, and at the same time reduce the calculation time and the storage to improve the efficiency. The planet-ring meshing element is a continuous dynamic meshing process. Hence, the dynamic load can be discretized into n impact loads for each period of meshing of the gear teeth. As shown in Fig.\u00a05a, the curve A is dispersed into 12 impact loads. From meshing in to meshing out, the action time of each impact load is \u0394t, and the action position successively changes from the first contact line entering into meshing to the exit of meshing. When meshing the ring gear, the traditional meshing method has dense mesh in the contact area, which will greatly increase the calculation time and storage. To avoid the above situation, when meshing the ring gear teeth, it is divided into n equal parts along the tooth profile, that is, n + 1 contact lines are divided, and the discrete load is applied to the corresponding contact lines, and then the 1 3 vibration and stress characteristics of the ring gear under the meshing excitation are obtained. When multiple teeth are meshed at the same time, there is also a problem of dynamic load distribution between teeth. In this paper, the LTCA method is used to discretize one period of the dynamic load on the ring gear, and the deformation of each participating gear along the contact line direction in a period is determined \u0394ij(t) (the deformation along the ith contact line of the jth engaged gear at t time), as shown in Fig.\u00a05. Fij(t) is the component force of the dynamic load on the jth engaged tooth along the ith contact line at time t, as shown in Fig.\u00a05b. And the meshing element of the planet-ring gear is single tooth and double tooth alternately meshing. According to the calculation Eq.\u00a0(5) of gear theoretical coincidence degree, in a meshing cycle, 93% of the time is in double tooth meshing, and 7% of the time is in single tooth meshing. The theoretical calculation formula of the coincidence degree of the planet-ring gear meshing element is where \u03b1a1, \u03b1a2 and \u03b1\u2032 are the planet, the ring gear tip pressure angle and the indexing circle pressure angle, respectively, z1 and z2 are the number of teeth of the planet gear and the ring gear",
            " To effectively extract the time-domain strain history of the gear root in the working process, this paper uses NI USB 6009 acquisition card (sampling frequency: 5000\u00a0Hz, system detection accuracy: 4.8\u00a0mV/\u03bc\u03b5). A sensor is pasted at the position of 1\u00a0mm at the root of the tooth space (number of sensors: 1), and then the time-domain strain history of the tooth root is obtained, as shown in Fig.\u00a09a. The dynamic strain time history can be divided into two parts by analyzing the above figure, that is, the meshing area and non-meshing area when teeth 6 and 7 (Fig.\u00a05b) participate in meshing. It can be seen from Fig.\u00a09 that in the meshing area, the meshing action leads to the complex meshing process from compression to tension at the tooth root, and 1 3 the dynamic strain will also appear sudden change. It can be seen from Fig.\u00a09a, b, the maximum dynamic strain measured by the experiment is 26.2\u00a0\u03bc\u03b5, and the maximum dynamic strain calculated by simulation is 28.5\u00a0\u03bc\u03b5. There is a certain error between the experimental measurement result and the simulation result, which is due to the numerical error caused by the test method and other factors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002542_icem49940.2020.9270965-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002542_icem49940.2020.9270965-Figure7-1.png",
        "caption": "Fig. 7. Modular stator and non-uniform tooth [26].",
        "texts": [
            " Innovative Topologies Some innovative topologies have been presented based on modular design, flux controllable and separated phase topologies. Meanwhile, novel motor topologies for fault tolerance from other industrial application are worthy to investigate, especially electric vehicles. 1) Modular Design The modular design philosophy could be achieved by modular winding, modular stator and multiple machine design. A modular stator PM synchronous machine has been proposed which applied for wheel-driving vehicle application as well [26], shown in Fig. 7. The 24 slots 14 poles combination, which exhibited the harmonics inhibition of magnetomotive force (MMF). The highlight of topology, concentrated winding configuration with unequal teeth widths, has completely isolation of thermal and magnetically. For specific aerospace actuator applications, two candidate segmentation strategies, axial segmentation and circumferential segmentation are adopted for fault tolerance and shown in Fig. 8. [23]. In the formal architecture, two separate stators with separate driven inverters are embedded into a common housing and rotor, adopting single layer FSCW configuration; the latter one, two set of three-phase in one stator with double-layer FSCW configuration driven by two independent inverters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003680_1999-01-0404-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003680_1999-01-0404-Figure1-1.png",
        "caption": "Figure 1. EHPAS System schematic",
        "texts": [
            " Performed durability tests verified the high level of reliability on the component subsystem and entire system level. Special attention was dedicated to investigate and optimize Magnetic Sensitivities, with the use of specialized Visteon test facility. The results of detailed DFMEA analyses have been incorporated in Diagnostic portion of the entire control strategy. The complete EHPAS system consist of the following components: DC motor, Pump, Transmission lines, Gear valve, and Actuator, as presented in Fig.1. Several complexities of the EHPAS system models have been developed and used for the analyses and control design. First, 2 complete non-linear model of the each separate component was developed and entire EHPAS system model was assembled with the use of component models. Second, linearization of the models has provided a useful tool for control design and analyses. The static model was developed based on a complete non-linear model and was also used for analyses. In this paper we present only non-linear model that is used for the final control algorithm optimization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001084_j.automatica.2019.108497-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001084_j.automatica.2019.108497-Figure3-1.png",
        "caption": "Fig. 3. The red surface is a desired trajectory X\u0302\u03a3 [0, T ] we want the ensemble formation system (9) to follow. Each X\u03c3 (t), for (t, \u03c3 ) \u2208 [0, T ]\u00d7\u03a3 , belongs to an open, path connected set Q . The red dashed curve is a desired trajectory for the individual system-\u03c3 . The gray surface is the trajectory X\u03a3 [0, T ] generated by a common control input u[0, T ] such that it is within the \u03f5-tubular neighborhood of X\u0302\u03a3 [0, T ]. The black solid curve is the trajectory X\u03c3 [0, T ] for the individual system-\u03c3 .",
        "texts": [
            " Outline of contribution. We establish in the paper a sufficient condition for the ensemble formation system (1) to be approximately path-controllable. Roughly speaking, this is about the capability of using a common control input to simultaneously steer every individual formation system in the ensemble to approximate any desired trajectory of formations. Note, in particular, that trajectories of different individual systems can be completely different. We refer to Definition 4 for a precise definition, to Fig. 3 for an illustration, and to Theorem 1 for the controllability result. The theorem addresses the interplay between the information flows within the individual formation systems (i.e., the common digraph G), the parameterization functions {\u03c1s} r s=1, and the controllability of the ensemble formation system. A key component of the analysis of the ensemble formation system involves computing the iterated Lie brackets of control vector fields of system (1), which further boils down to the computation of iterated matrix commutators of certain sparse zero-row-sum matrices",
            " System (9) is approximately ensemble path-controllable over Q if for any smooth target trajectory X\u0302\u03a3 [0, T ], with X\u0302\u03c3 (t) \u2208 Q for all (t, \u03c3 ) \u2208 [0, T ] \u00d7 \u03a3 , and any error tolerance \u03f5, there are integrable functions uij,s : [0, T ] \u2192 R as control inputs such that the trajectory X\u03a3 [0, T ] generated by (9), from an initial condition X\u03a3 (0) with X\u03c3 (0) \u2208 Q and \u2225X\u03c3 (0) \u2212 X\u0302\u03c3 (0)\u2225 < \u03f5 for all \u03c3 \u2208 \u03a3 , satisfies \u2225X\u03c3 (t) \u2212 X\u0302\u03c3 (t)\u2225 < \u03f5, \u2200(t, \u03c3 ) \u2208 [0, T ] \u00d7 \u03a3 . We illustrate the above definition in Fig. 3: With the above preliminaries, we are now in a position to state the first main result of the paper (compared to Lemma 1): Theorem 1. Let G be strongly connected and N > (n + 1). Suppose that the set of parameterization functions {\u03c1s} r s=1 separates points and contains an everywhere nonzero function; then, system (9) is approximately ensemble path-controllable over the set Q of nondegenerate configurations. A sketch of proof will be given at the end of the section. Detailed analysis will be provided in Sections 4 and 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001759_j.mechmachtheory.2020.103995-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001759_j.mechmachtheory.2020.103995-Figure6-1.png",
        "caption": "Fig. 6. The kinematic paths of \u03b83 vs \u03b81 for linkage with a = 15 mm, r 1 = 20 mm, r 2 = 40 mm and \u03b1= 50 \u00b0, where (a)-(c)-(d)-(f) correspond to bifurcation configurations. (b) and (e) are configurations along type I and type II.",
        "texts": [
            " Its quick-return ratio Kis K = 2 OA 2 OB = \u03c0 + 2 arcsin ( a ( r 2 \u2212r 1 ) tan \u03b1 ) \u03c0 \u2212 2 arcsin ( a ( r 2 \u2212r 1 ) tan \u03b1 ) . (29) Note that we can also get a double rocker linkage if link 23 in Fig. (5) is selected as the frame. But the case of double rocker will not be discussed further because it has the same design parameters as crank-rocker linkage. Next, we shall discuss the bifurcation behaviours corresponding to the above crank-rocker and double crank linkages. For the crank-rocker linkage, | a ( r 2 \u2212r 1 ) tan \u03b1 | < 1 , as demonstrated in the Fig. 6 with design parameters a = 15 mm, r 1 = 20 mm, r 2 = 40 mm and \u03b1= 50 \u00b0 as an example, there are only two bifurcation points, B 1 (- \u03c0 /2, 0) and B 2 ( \u03c0 /2, 0), among the kinematic paths of type I (red dotted curve), type II (black solid curve) and open-chain branch (green dotted line). At each bifurcation point, type I and type II linkages are in the different configurations, which means that type I can never be di- rectly bifurcated into type II. Hence, there are bifurcations between type I and open-chain branch or type II and open-chain branch"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000699_j.ifacol.2016.07.974-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000699_j.ifacol.2016.07.974-Figure1-1.png",
        "caption": "Fig. 1. Schematics of the quadrotor.",
        "texts": [
            " Eindhoven, The Netherlands Copyright \u00a9 2016 IFAC 1 282 Stanislav Tomashevich et al. / IFAC-PapersOnLine 49-13 (2016) 281\u2013286 Centre Ob is the center of mass quadrotors. The main objective is to move quadrotors to the desired point in the IRF, while the movement is accompanied by a tilting mechanism that occur in the IRF and NRF in different ways, so it is needed to properly observe transitions from one system to another. The angles at which quadrotor deflected relate to a fixed system are the Euler angles \u03b3, \u03d1 and \u03c8. Figure 1 presents two systems of reference, the main directions and angles that form the overall motion of the system. Linear quadrotors position is described by the variables x, y, z \u2013 coordinates of its center of mass. Quadrotor has a symmetrical structure with four engines arranged on beam axes ObXb and ObZb at a distance l from the center of mass. The matrix of inertia is as follows: IA = [ Ix 0 0 0 Iz 0 0 0 Iy ] , (1) where Ix, Iy, Iz are rotation inertia moments with respect to the corresponding axes, Ix = Iz",
            " Eindhoven, The Netherlands Stanislav Tomashevich et al. / IFAC-PapersOnLine 49-13 (2016) 281\u2013286 283 Centre Ob is the center of mass quadrotors. The main objective is to move quadrotors to the desired point in the IRF, while the movement is accompanied by a tilting mechanism that occur in the IRF and NRF in different ways, so it is needed to properly observe transitions from one system to another. The angles at which quadrotor deflected relate to a fixed system are the Euler angles \u03b3, \u03d1 and \u03c8. Figure 1 presents two systems of reference, the main directions and angles that form the overall motion of the system. Linear quadrotors position is described by the variables x, y, z \u2013 coordinates of its center of mass. Zb Xb Yb Ob Z X Y O N \u03d1 Zm Xm Ym R4 R2 R1 R3 \u03c8 \u03b3 \u03c91 \u03c92 \u03c93 \u03c94 Fig. 1. Schematics of the quadrotor. 2.2 Quadrotor design Quadrotor has a symmetrical structure with four engines arranged on beam axes ObXb and ObZb at a distance l from the center of mass. The matrix of inertia is as follows: IA = [ Ix 0 0 0 Iz 0 0 0 Iy ] , (1) where Ix, Iy, Iz are rotation inertia moments with respect to the corresponding axes, Ix = Iz. Quadrotor can be represented as a sphere attached to the ends of the beams with four other spheres, which mean the engines. The model also takes into account the inertia of screws provided in the form of bars fixed in the middle on an axis perpendicular to the screws"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000183_tim.2019.2949319-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000183_tim.2019.2949319-Figure8-1.png",
        "caption": "Fig. 8. DCW SWS with bends, sections with tapered pillar heights, and step-tapered waveguide to connect to WR-28 flanges.",
        "texts": [
            " The linear coupler DCW SWS includes two regions, in the range of 24% of overall length, where the electron beam is partially modulated and only a very weak interaction is present. However, the electron beam traveling in these regions still needs magnetic focusing. In other words, the regions increase the volume without adding to the gain. A novel approach, based on a 90\u00b0 bending of the DCW at the ends of the main interaction region, has been developed. This places the regions of weak interaction, with tapered pillars, orthogonal to the main gain section (Fig. 8). This approach permits a substantial reduction of the overall TWT length. Fig. 8 shows the complete optimized millimeter wave structure. The DCW is bent at 90\u00b0 with respect to the beam tunnel and the collector end. After each bend follows a DCW to rectangular waveguide transition, to couple the hybrid mode to the TE10 mode. Twenty pairs of DCW pillars are tapered in height and in period length in the DCW transition. The DCW enclosure width has a sinusoidal profile. The total length of the DCW transition is 69 mm. Finally, a matching section connect the DCW transition to the WR-28 flange"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003112_s12555-020-0541-3-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003112_s12555-020-0541-3-Figure2-1.png",
        "caption": "Fig. 2. Structural sketch and side view of the parallel robot for AECC.",
        "texts": [
            " For this kind of general nonlinear systems with stochastic disturbances [35,36], the stochastic disturbances can be incorporated into the lumped disturbance term, thereby the proposed FRSMDC based on disturbance estimation technique can also be effective. To verify the validity of the traditional SMC law (6) and the proposed FRSMDC law (11) for parallel robots, a 6-DOF parallel robot with a bilateral symmetrical structure for AECC, is studied as an example for simulation. The structural sketch and side view of the 6-DOF parallel robot are illustrated in Fig. 2. To indicate that the proposed controller can be adapted to parallel robots with both translational and rotational degrees of freedom, the active joints and the end effector of the parallel robot used for case study can perform both translational and rotational motion. Specifically, there are six active joints, i.e., four sliders responsible for translational motion and two driving wheels responsible for rotational motion. Besides, as shown in the first picture in Fig. 2, the end-effector of the parallel robot can perform not only translational motion in Z-direction, but also rotational motion around Y -axis. The physical parameters of the parallel robot for AECC are listed in Table 1. The friction term [7] can be expressed as N(t) = F c sgn(q\u0307)+Bcq\u0307, where F c = diag(1.1, 1.1, 1.1, 1.1, 1.1, 1.1) is the Coulomb friction and Bc = diag(1.5, 1.5, 1.5, 1.5, 1.5, 1.5) is the coefficient of the viscous friction. To verify the robustness of the proposed method, the external disturbance is set to \u03c4dis(i) = 3sin(2\u03c0t +\u03c0/2)(N \u00b7m) (i = 1, 2, \u00b7 \u00b7 \u00b7 , n) and the modeling error is set to \u2206M = diag(1, 1, 1, 1, 1, 1)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure12-1.png",
        "caption": "Fig. 12. The meshing force in 3D conditions with two springs, one on the top side, one on the bottom side (a) out-of-penetration condition (b) the compression of the springs under a non-even load to its top and bottom side.",
        "texts": [
            " However, one additional force Fc2 needs to be introduced to prevent an unrealistic gear position, where two gears as rigid bodies are penetrating each other. On each transverse plane of the gear pump, the pitch diameter 2rp is the minimum rigid-body center distance between two gears, below which the transverse profile to two gears will overlap, which indicates a physical penetration. When such a penetrating gear position is given by the gear position model, a reaction force resulted from solid contact will be triggered to push two gears in opposite directions against each other to overcome the extra external load, as shown in Fig. 12. This additional contact force in this simulation model is modeled as two springs, connecting two gears on top and bottom ends, respectively. When the springs are triggered, their force will be in the direction along the centerline of two gear shafts, pushing two gears away from each other, as shown in Fig. 12. The condition triggering up this force is that the center distance predicted by the journal bearing model is smaller than the pitch diameter. Based on the assumption that this contact force is in the elasticity regime, the magnitude of the contact force is proportional to the difference between the instantaneous center distance and the pitch diameter, i.e. Fc2 = { K ( 2rp \u2212 i ) for i \u2264 2rp 0 for i > 2rp (44) X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 where K is the meshing stiffness of the gear, which is assumed to be Young\u2019s modulus of the material that the gears are made of. This force is zero when the current center distance is greater than the pitch diameter, which means the gear is in an off-contact condition. The condition Eq. (44) is evaluated for the top and bottom end of the gears separately. For example, in a misaligned case represented by Fig. 12b, the penetrating condition only happens on the top end, therefore only the spring on the top is triggered and the force will be applied on two gears on the plane z = H/2 with opposite directions. For 3D gear dynamics in a CCHGP unit, there are four possible modes of micro-motions. The first mode is the transverse micromotion, for which the shafts of two gears remain parallel to each other, and the centers of the gear shafts shift in the transverse plane, as shown in Fig. 13a. The second mode is the misalignment that gives a non-parallel position of two gear shafts (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001825_tcst.2020.3009636-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001825_tcst.2020.3009636-Figure2-1.png",
        "caption": "Fig. 2. (a) Visualization and labeling of the 1-D (vertical-only) jumping robot model. (b) Illustration of the jumping task; the robot must thrust its motor mass to jump off the ground and achieve a specified jump height at its apex.",
        "texts": [
            " The robot quickly accomplished specified control tasks, each within 10\u201315 control-learning iterations. Using this model-based data-driven learning and control approach, each terrain model was learned, and specified tasks were completed precisely, while only requiring a small set of experiential data. This section reviews the equations governing 1-D jumping, as applicable to any surface type, rigid or deformable. The equations are arranged to highlight terrain substrate forcing, Fsub, the unknown quantity in this system. The 1-D jumping model, shown in Fig. 2 and presented in [16], is modeled by three massed bodies: a linear motor, a thrust rod, and a foot. This robot model jumps by applying force between the motor and thrust rod, which can be leveraged to pump energy into the spring-loaded system. Vertical motor actuation drives rod displacement which, in turn, transmits forcing to the foot through the connecting spring (which is mass-less, but modeled with linear viscous damping). The closed-form system dynamics are x\u0308 f = \u2212g + 1 m f [k(\u03b1 \u2212 \u03b1\u0304) + c\u03b1\u0307] + 1 m f Fsub \u03b1\u0308 = \u2212 [ m f + mr + mm m f (mr + mm) (k(\u03b1 \u2212 \u03b1\u0304) + c\u03b1\u0307) + mm mr + mm \u03b2\u0308 ] \u2212 1 m f Fsub (1) where body masses of the motor, rod, and foot are designated mm , mr , and m f , respectively",
            " Each discrete point, i , is rendered dynamically consistent with the next point, i + 1, via \u201cdefect\u201d constraints in the NLP, which approximate implicit integration ( xd i+1 \u2212 xd i ) \u2212 1 2 ( td i+1 \u2212 td i )( x\u0307d i+1 + x\u0307d i ) = 0 (11) x\u0307d i \u2212 f ( td i , xd i , ud i ) = 0 (12) for all i , derived from a trapezoidal integration scheme. The equations of motion f include the GP-based forcing model Fsub in symbolic form. The control task is to achieve a specified \u201cjump height\u201d defined as the difference between the initial rod height, x0 r , and the rod\u2019s highest point during the jump, x f r , as shown in Fig. 2(b). Consequently, targeted jumping requires adding an equality constraint x f r \u2212x0 r = h\u2217 with h\u2217 equal to the target jump height. In hardware implementation, the true actuation signal, \u03b2\u0308, arises from a low-level trajectory tracking control loop applied to the linear motor of Fig. 2(a). The closed-loop motor dynamics introduce a dynamic response that impacts jumping performance. We explicitly account for these tracking dynamics in the optimization by modeling the control loop as part of the system dynamics [4]. In brief, the solution, u, is computed with anticipation, both of how the tracking will perform given the closed-loop dynamics and the saturation limits of the actuator force, Fm . The objective function is defined J(w) = \u2211 d\u2208D [ td 2 \u2212 td 1 2 J ( xd 1 , x\u0307d 1 , ud 1, \u03b2\u0307 \u2217,d 1 ) + N\u22121\u2211 i=2 td i+1 \u2212 td i\u22121 2 J ( xd i , x\u0307d i , ud i , \u03b2\u0307 \u2217,d i ) + td N \u2212 td N\u22121 2 J ( xd N , x\u0307d N , ud N , \u03b2\u0307\u2217,d N )] (13) where J (\u00b7) is the integrand of a continuous-time cost",
            " The foot of the robot is the only component to interact with any jumping surface and is composed of a lightweight, 3-D printed cylinder with base radius, 38.1 mm, and height, 35 mm. Fig. 4(b)\u2013(d) shows the three categorically distinct surface types employed during the experiments: solid ground, a trampoline surface, and a GM bed. Performance of Algorithm 1 was evaluated for control tasks requiring the hopper to achieve a range of targeted peak jump heights. Jump heights were defined to be the maximum tracked height attained by the rod marker, relative to its resting position at the start of the jump, shown in Fig. 2(b). To facilitate the measurement of state variables x f and xr , spherical, white markers were mounted to both the foot and rod, respectively. A gray-scale PointGrey Grasshopper camera with a high-frame rate (200 Hz) was positioned to simultaneously capture both markers during the course of each experiment. Marker tracking was performed frame-to-frame. In the interest of repeatability and consistency across experimental trials, we defined and followed an experimental protocol, with minor adjustments for each terrain tested"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002247_tmech.2020.2972295-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002247_tmech.2020.2972295-Figure15-1.png",
        "caption": "FIGURE 15. UGSIGN posterior separation plotted versus prior LPR for various values of DOF count at GSNRs of 10 dB and 20 dB.",
        "texts": [
            " (117) Equation (114) is plotted versus prior LPR for selected values of GSNR in Figure 13 at DOF counts of 1, 4, and 16; some further detail of the near zero LPR region, together with the null statistic lower bound, is provided in Figure 14. In these plots, the dashed red line denotes the boundary 3\u2212 = M ln (1+ ESNR). In Figures 13 and 14, the mathematical requirement that posterior separation increase with GSNR for constant prior LPR is easily seen. However, note that there is no requirement that the converse be true; that is, posterior separation is not required to increase with prior LPR at constant GSNR (although that is typically the case). In Figure 15, the results of Figure 13 have been regrouped to now show posterior separation for different DOFs at constant GSNR. As should be expected from the earlier density results, Figure 15 demonstrates graphically that KSIGN results again emerge as DOF counts grow large. However, even for finite DOF counts, these results are remarkably similar to the comparable KSIGN results seen in Figure 4, demonstrating the rather mild dependence of information measures on density choice. The principal 1474 VOLUME 2, 2014 difference appears to be an adjustment for asymmetry between negative and positive LPRs, the fundamental source of which is the asymmetric nature of the densities involved. It is interesting to note the effect of increasing DOF count in Figure 15; when prior LPR is negative (corresponding to the typical rare target situation), additional DOFs (for the same GSNR) improve performance, whereas when prior LPR is positive (i.e., the atypical target rich situation), the reverse is true. The relative independence of information measures on density choice is further reinforced by the limiting forms of UGSIGN posterior separation. Because GSNR = MESNR2 1+ ESNR , (118) the large observation limit can be achieved one of two ways. As previously mentioned, in the large DOF case, the KSIGN results are recovered exactly, so that SEP+LO(D) \u223c GSNR/2+ ( P\u22121 \u2212 P \u2212 0 ) 3\u2212, (119) while in the large ESNR case SEP+LO(E) \u223c P\u22121 GSNR+ ( P\u22121 \u2212 P \u2212 0 ) 3\u2212"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001627_icrom48714.2019.9071917-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001627_icrom48714.2019.9071917-Figure3-1.png",
        "caption": "Fig. 3: The effect of the horizontal speed on the blade disk which is the cause of blade flapping [10].",
        "texts": [
            " In this section, the most blade flapping, Vortex Ring State (VRS) and drag aerodynamic effects are described. 1) Blade Flapping: When a quadcopter has a horizontal speed, Blade flapping effect is occurred. The blade that has speed in the direction of drone\u2019s horizontal speed, has more relative speed in comparison of the blade that has a speed in opposite direction (in +\u03c0 phase). By assuming this difference in two sides of the blade disk, there is a torque due to thrust difference produced by the two sides of the blade disk Fig. 3 and the blade flapping is occurred. This point should be taken into consideration that there is a hinge in helicopter blades, but in quadcopter blades, there is not any degree of freedom in flapping direction and the effect acted on the blades by deflecting them. For detailed formulation of blade flapping effect, see [10]. 2) Vortex Ring effect: In the descent phase of the quadcopter, depending on the drone velocity, the quadcopter experiences high fluctuation on its blades. Sometimes these fluctuations are the cause of a crash during the descent maneuvers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002572_j.apm.2020.12.020-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002572_j.apm.2020.12.020-Figure7-1.png",
        "caption": "Fig. 7. First three orders of vibratory mode under Pose II.",
        "texts": [
            " Since the number of flexible beam elements in ANSYS software is far more than that used in the dynamic model of this paper, the simulation deviations are acceptable. The vibratory mode schematic diagrams of the Table 2 Frequencies under Pose I (Unit: Hz). Order No. First order Second order Third order Theoretical elasto-dynamics model 9.279 9.282 29.622 ANSYS software simulation 9.829 10.138 33.332 Table 3 Frequencies under Pose II (Unit: Hz). Order No. First order Second order Third order Theoretical elasto-dynamics model 7.188 10.352 28.595 ANSYS software simulation 7.866 9.342 31.622 first three natural frequencies are shown in Fig. 6 to Fig. 7 , in which the first order of vibratory mode is the mechanism vibrating left and right along the y axis, the second order of vibratory mode is the mechanism vibrating back and forth along the x axis, and the third order of vibratory mode is the torsional vibration of the mechanism around the z axis. In general, the Lagrangian multiplier must be solved simultaneously when solving the system dynamic equations of index-1, which inevitably increases the complexity of the solving process. In this section, we propose a dynamic equation solving strategy which combines Baumgarte stabilization technique and Gill algorithm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003630_047134608x.w1111-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003630_047134608x.w1111-Figure2-1.png",
        "caption": "Fig. 2. Control-system elements in a modern piloted aircraft.",
        "texts": [
            " In the early days of flight, this was accomplished solely by the pilot, who determined the desired trajectory and speed and, through training, moved the control surfaces by means of hand and foot controls (referred to as cockpit inceptors) so as to produce the desired results. In simplest terms, the pilot commanded the desired aircraft trajectory and speed, sensed the actual aircraft trajectory, and actuated the control surfaces in response to the differences in these quantities. This process constitutes the activity of a feedback control system. As shown in Fig. 2, much for the responsibility for feedback in the case of modern high-performance aircraft has been assumed by inanimate systems, incorporating accurate sensors and precise, powered actuators to move control effectors in a manner prescribed by a flight control law typically implemented on an onboard digital computer, frequently referred to as the flight control computer. Indeed, in most modern high-performance aircraft, the pilot is no longer directly connected to the control effectors by means of cables, pulleys, etc",
            " Since the bandwidth requirements for the h and u loops are identical, either may be chosen first. Here, the h loop was selected, with the resulting Gh shown in Table 2. Figure 10 shows the resulting effective vehicle, with hc and \u03b4T serving as inputs. Finally, the airspeed loop is closed with the Gu shown in Table 2. The final effective vehicle is shown in Fig. 11, now with hc and uc serving as inputs. Digital Implementation of the Control Law. The compensator transfer functions G\u03b8(s), (s), Gh(s), and Gu(s) of Table 2 constitute the flight control law for this application. As Fig. 2 indicates, such laws are now routinely implemented in a digital flight control computer. Thus, the control law must be expressed in a form suitable for digital implementation. To accomplish this, the nested control architecture of Fig. 6 is redrawn in an equivalent form where the compensators act in parallel. This can be easily done by referring to Fig. 6 and writing expression for the control inputs as Figure 12 shows the resulting control structure. The A/D and D/A blocks refer to analog-to-digital and digital-to-analog converters, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003006_tmag.2021.3085509-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003006_tmag.2021.3085509-Figure1-1.png",
        "caption": "Fig. 1. 3D views of (a) Radially-Magnetized TROMAG (RMT), and (b) Reluctance Radially-Magnetized TROMAG (R-RMT). 2D views of (c) RRMT with rectangular iron teeth, (d) Design region of the topology optimization model, and (e) R-RMT and its NGnet circles.",
        "texts": [
            " HMT Halbach-magnetized TROMAG (with PMs on both components). R-RMT Reluctance radially-magnetized TROMAG (with PMs on one component and variable reluctance structure on the other) R-HMT Reluctance Halbach-magnetized TROMAG (with PMs on one component and variable reluctance structure on the other) I. INTRODUCTION ROMAG is a magnetic device that converts high-force, low-speed linear motion to low-torque, high-speed rotation, and vice versa [1]. The device is also referred to as magnetic screw [2] and magnetic lead screw [3]. As shown in Fig. 1a, a TROMAG consists of an inner and an outer component which are separated by an air gap. One component serves as the rotor and the other as the translator. The linear to rotary motion conversion is achieved via helically-shaped permanent magnet (PM) poles. TROMAG can offer large values of shear stress and gear ratio; resulting in light-weight and compact designs. In addition, similar to other types of magnetic gears, TROMAG offers low noise and high reliability. Advantages of TROMAG have made it suitable for a wide range of high force, low speed linear motion applications. Helical PM poles can assume different magnetization patterns. Radially-magnetized TROMAG (RMT), axiallymagnetized TROMAG (AMT), and Halbach-magnetized TROMAG (HMT) have been proposed. Another configuration, shown in Fig. 1b, is the reluctance TROMAG in which one component is furnished by helical PM poles while the other component has a variable reluctance structure. The reluctance TROMAG offers less shear stress compared to configurations having PMs on both components. However, lack of PMs on the variable reluctance component makes this configuration potentially advantageous in, for example, applications with a long stroke. Reluctance TROMAG is investigated in multiple works; with radially-magnetized PMs (R-RMT) in [2], with Halbachmagnetized PMs (R-HMT) in [3], and with axially-magnetized PMs (R-AMT) in [4]",
            " Here, the on/off method is applied to Gaussian circles mapped to the design region. The method is incorporated in JMAG-Designer software, which is employed in this paper for optimizing the reluctance TROMAG. Part II presents topology optimization of the reluctance TROMAG for two magnet configurations; R-RMT, and RHMT. In part III, novel topologies obtained in part II are further investigated using parametric analysis, and compared against the conventional configurations. Conclusions are summarized in part IV. Fig. 1c shows 2D view of a conventional R-RMT where the reluctance component has iron core and rectangular teeth. The goal is to find reluctance component topologies, if any, that outperform the conventional one. To that end, the reluctance component is assumed to be the design region of topology optimization model, and topology optimization coupled with GA is applied to two reluctance TROMAG configurations; RRMT and R-HMT. To perform topology optimization, the design region is divided into quadrilateral elements (Fig. 1d), where each element could be either air or iron. Note that if quadrilateral elements in a resultant topology are not physically connected, the topology is impractical. To avoid such cases, a constraint on elements connection is applied to the optimization process. NGnet circles are shown in Fig. 1e. From experience, in order for the GA to converge, the number of circles is set to 21. A larger number would result in longer computation time without noticeable improvement in results. In all cases, steel 1020 is used for ferromagnetic iron core and N52 grade NdFeB is used for magnets. In practice, it makes sense to use the reluctance component as the long part, and the PM component as the short part. However, in simulations the reluctance component is set as the shorter part so as to save on computation time",
            " For each individual produced by the GA, the reluctance component is moved by one pole pitch while the PM component is held stationary. This way, one complete cycle of the force characteristic is obtained and the individual\u2019s shear stress is calculated. With the population size of 67, the GA converged to Fig. 2a after ~38000 cases. Note that the core material for the reluctance component of Fig. 2a is non-ferromagnetic. Another well performing topology produced by the GA is shown in Fig. 2b. Compared to the conventional R-RMT with optimized tooth dimensions, (Fig. 1c), topologies shown in Fig. 2a and Fig. 2b offer 40% and 10% higher shear stress, respectively. Jagged corners of the new R-RMT topologies are smoothed manually, resulting in Figs. 2c and 2d. The topology shown in Fig. 2c is referred to as the R-RMT with elliptical teeth and air core and will be further investigated in part III. Noting that the topology shown in Fig 2d shows only 10% improvement over the conventional configuration, it is not studied any further in this paper. In a quasi-Halbach configuration with two segments per pole, the ratio of radially-magnetized segment width to the pole pitch is referred to as the radial magnetization percentage (RMP), and is an aspect of HMT design optimization",
            " more shear stress compared to the optimized conventional RHMT. Repeating the process for the conventional R-HMT with RMP of 75% resulted in elliptical iron teeth with air core, shown in Fig. 3c. This topology offers almost the same shear stress as the optimized conventional R-HMT. The second best case obtained for the case of 75% RMP is shown in Fig. 3d, and it offers 10% less shear stress compared to the optimized conventional R-HMT. In the topology optimization presented in the previous part, the PM thickness was fixed at 5 mm and the element size (Fig. 1d) was set to 0.5 mm. In this section, the optimum topologies obtained in the previous part are further investigated using a parametric analysis. As seen in Part II, the optimum topology for the reluctance component of the R-RMT is elliptical teeth and air core, offering 40% more shear stress compared to the conventional R-RMT. Note that since this topology uses non-ferromagnetic material as the core of the reluctance component, it would result in additional savings on the overall weight as well. The 3D and 2D models of this topology along with flux lines at the pull-out force position are shown in Figs",
            " Compared to the elliptical teeth with air core, the rectangular teeth with air core topology is more sensitive to an increase in its tooth height. For the optimum case of Fig. 7, increasing H form 1.4 mm to 3 mm reduces the shear stress to 71 kN/m2. In comparison, the elliptical iron teeth with air core and rectangular iron teeth with air core respectively produce 40% and 29% more shear stress than the conventional R-RMT. Fig. 8a compares the shear stress values of the three R-RMT topologies (Figs. 4b, c, and d) with that of the RMT (having PMs on both the rotor and translator components; Fig. 1a). Clearly, all reluctance RMT topologies offer less shear stress compared to the RMT. In order to evaluate PM utilization, shear stress over normalized magnet volume is calculated. The base value of magnet volume is set to the volume of magnet in the short component of the RMT regardless of magnet thickness. Figs. 8b, c, and d show the shear stress over normalized magnet volume when the long component is respectively 2, 3, and 4 times longer than the short component. It is observed that as the difference between the lengths of the short and long components increases, R-RMT eventually surpasses the RMT in terms of magnet utilization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001834_j.addma.2020.101547-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001834_j.addma.2020.101547-Figure3-1.png",
        "caption": "Fig. 3. Coherent Gradient Sensing system.",
        "texts": [
            " 2, the collimated laser beam carrying the deformation information is reflected from the specimen surface and reflected light is split into two identical beams of incident light with the help of a beam splitter combination. The two incident light beams pass through two different shearing devices. The lines of grating in the first and second shearing devices are parallel to the x- and y-axes, respectively. According to the experimental set-up for improved-CGS schematic, a CGS system was designed for in-situ monitoring of the DED process in this study, as shown in Fig. 3. The laser spot was turned into a parallel light beam using a collimation system. With the help of a filtering aperture, the interference fringes pattern was obtained by camera. In this study, the grating pitch was p=25 \u03bcm, the distance between G1-2 and G2-2 was \u0394 =38mm, the distance between G1-1 and G2-1 was \u0394 =38mm and the laser wavelength was \u03bb=532 nm. In this study, a diode laser with a wavelength of 1064 nm was used as the energy source and argon shielding gas was applied to protect the surfaces of product in the DED process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001811_s12008-020-00670-z-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001811_s12008-020-00670-z-Figure9-1.png",
        "caption": "Fig. 9 Assigning the coordinate frames for the links of the robot: a the assignment of the coordinate frames is correct; b one of the coordinate frames is incorrect",
        "texts": [
            " For example, the professor explained the concept of tool-center point (TCP), how to define it on a manipulator, and showed the consequences of an incorrect definition (i.e., the difference between defining the TCP at the bottom and defining the TCP at the top in a pick task). The use of computer simulations allowed students to analyze the consequences of changing the position of the TCP and the end-effector (the differences among the definitions of the end-effector at different points of the manipulator); consequently, they learned the right way to define these conditions. Figure\u00a09 shows how the professor explained that the choices of the various coordinate frames are not unique and showed examples of correct and incorrect assignments. In Fig.\u00a09.b, it can be seen that the reference system enclosed in the red circle is separated from the robot; the cause is that the correct parameters were not defined in the previous reference frame. Sometimes, these errors arose when the students simulated various movements of the robot. Another typical error is that students used the wrong units to set the translation operations; that is, meters were selected instead of centimeters for these operations. 3. Solving exercises in the virtual environment: The professor guided the students to solve a set of exercises (manipulators of 2-DoF, 3-DoF, and 4-DoF with prismatic and revolute joints) about forward kinematics using computer simulations. The problems were about the relationship between the joints of the manipulator and the orientation and position of the end-effector. Using computer simulations, the students verify stepby-step their results in order to identify and correct their mistakes (i.e., the definition of the reference system, the calculation of the DH parameters, and the values of the homogeneous transformation matrix). For example, Fig.\u00a09 shows the simulation screen of a robot; the students run the simulation with the DH parameters that they previously calculated using the DH method. If the joint movements and/or the position of the end-effector are not correct, students must review their procedure to calculate a new set of DH parameters. The main advantage of the use of computer simulations in comparison with a traditional teaching method is that students can verify their results because the reference system of each joint and the TCP of the robot are updated according to the change of position of each joint"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000222_0954406219885979-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000222_0954406219885979-Figure8-1.png",
        "caption": "Figure 8. Force analysis of the double-row self-aligning ball bearing.",
        "texts": [
            " In details, when y is positive, the diagonal stiffness coefficient kxx decreases with y, while kyy and kzz increase with y; when y is negative, kxx, kyy and kzz decrease with y; kxx and kyy are almost unchanged but kzz increases with x. Meanwhile, the distribution of the cross stiffness coefficients kxy and kyx with respect to the tilting angles x and y exhibits a saddle shape; kxz and kzx increase with y, while kyz and kzy increase with x. In the case study, the diagonal stiffness coefficients kxx, kyy, and kzz are more sensitive to y than x. This can be explained by a force analysis of the double-row self-aligning ball bearing, as illustrated in Figure 8. The external force Fx applies on the inner ring, only a few balls at the top are compressed due to the clearance. It can be seen from Figure 8 that, when the inner ring tilts around y-axis, the contact angles of the compressed balls are approximately equal to the linear superposition of the initial contact angle ( 0\u00bc 12.7 ) and the tilting angle y; however, when the inner ring tilts around x-axis, the contact angles of these balls almost keep constant. From equations (18) to (21), the bearing stiffness is closely related to the contact angle, so kxx, kyy, and kzz are more sensitive to y than x. Figure 9 shows the contact force distribution of the double-row self-aligning ball bearing with three tilting angles"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000212_mees.2019.8896501-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000212_mees.2019.8896501-Figure1-1.png",
        "caption": "Fig. 1. Models of the axial magnetic bearing: geometric (left) and finite element (right) ones.",
        "texts": [
            " A finite element problem formulation is: }U{}U{}U{ }L{}L{}U]{K[ 1 res EMEMEM kk k , 3 where [KEM] is an electromagnetic problem coefficient matrix, {U} is a nodal potential vector, { U} is a potential vector increment, {LEM} is an applied load vector (electric current, voltage or magnetic induction), {LEM res} is a residual load vector, k is an iteration number. According to the Maxwell's stress tensor (T\u0305M), the magnetic force acting on a surface S, limiting a volume v, can be calculated by the following formula: S aa S MM dSHd ] 2 1 )([S 2 nnHHTF , 4 where n\u0305 is a unit vector, normal to the surface. Geometrical and finite element models of the axial AMB are shown in Fig. 1. The following notation is entered: 1 is a rotor made of a non-magnetic material, 2 is an axial AMB disk made of a ferromagnetic material located on the rotor, 3 and 4 are axial AMB stators with windings 5. The center of a fixed Cartesian coordinate system is located in a center of gravity of the AMB, which disk is in a central position. A nominal axial clearance on both sides of the disk is a. The stators of the axial AMB are made of steel, each with internal and external poles. There are coils with currents in the conductors of these coils between the poles of the AMB stators",
            " Let be the currents in the windings of the left and right coils denoted by ic1 and ic2 (from the positive and negative directions of the Oz axis). The voltages uc1 and uc2 (active electrical resistance windings rc1 and rc2) are applied to the windings of each pole coil to create magnetic forces and keep the disk near the center position. The rotor in the AMBs as a control object is unstable. In order to stabilize it and give to the suspension the necessary dynamic parameters, it uses a control system (CS) [4]. We introduce the control law with respect to the coordinate system in Fig. 1. Simplistically, it can be represented as dependencies of average currents in the windings of electromagnets on the rotor displacement [5, 6]: minminmax2,1 /)( iziii ac , 5 where the displacement currents imin in both windings at the central position of the rotor create a force tension and provide the required AMB stiffness. The technique of numerical experiments is as follows. To calculate the dependences of the magnetic forces on the rotor displacement, the full nominal gap (on both sides of the disk) is evenly divided into 2n+1 levels, so that the (n+1)th level coincides with the central position of the disk between the stators",
            " CALCULATION OF POWER CHARACTERISTICS OF AXIAL As examples of a practical use of the methodology, the analysis of two AMBs was considered. The first one is an axial AMB of a laboratory unit, which implements a passively active full magnetic suspension of a rotor [5, 6] (for a verification of the technique). The second is a radial classical eight-pole AMB (as an example of a possibility to analyze an AMB of a complex configuration). A numerical determination of power and stiffness characteristics was performed for an axial AMB with the following parameters (Fig. 1): outer and inner diameters of electromagnets (EM) are 0.12 and 0.05 m, a length of each EM in the axial direction is 0.021 m, outer and inner diameters of windings are 0.1 and 0.072 m, a number of turns is 300, winding resistance is rc=5 Ohm, a disk thickness is 0.02 m, a nominal clearance on each side is a=0.003 m. Fig. 2 shows the results of electromagnetic analysis for two positions of the rotor \u2013 when the rotor is displaced in the axial direction by -2 a/3 and at the central position of the disk in the AMB gap"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001564_j.robot.2020.103482-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001564_j.robot.2020.103482-Figure2-1.png",
        "caption": "Fig. 2. On the left: Torus model with wheel plane, wheel axis, and groundwheel frame axes, and a contact point placement in sphere and torus models for a non-zero camber angle. On the right: Photo of the CENTAURO robot wheel.",
        "texts": [
            " The former is an extrinsic parameter and no specific properties can be presumed; however, for the sake of simplicity, it is assumed that the ground normal is constant in the proximity of a contact point. The latter, as an intrinsic characteristic, can be modelled with more details. A common simplification is to consider the wheel as a sphere or a cylinder that provides a sufficient approximation for systems with fixed camber angle. However, the cylindrical geometry does not allow to model a non-zero camber angle without taking into account deformations, while the spherical geometry implies the wheel centre and the contact point are collinear along the ground normal, see Fig. 2. In this work, we propose to model the wheel as a torus to more closely represent the kinematics of the groundwheel contact for a non-zero camber angle. Even though the zero camber angle assumption is common in the modelling of the hybrid legged-wheeled platforms, when this motion is permissible by the robot mechanical design, maintaining the zero camber angle throughout the motion cannot be safely ensured when the robot is moving on uneven terrains. When the robot adopts a non-zero camber angle position while passing through a difficult terrain, the motion controllers developed with the zero-camber angle assumption cannot ensure the stable motion. Furthermore, consideration of the non-zero camber angle position allows the robot to extend its leg workspace with comparison to the standard zero camber angle controllers. Particularly, the proposed model is compatible with the geometry of the wheel used in the CENTAURO robot given in Fig. 2. The top photograph of Fig. 3 shows the CENTAURO robot in a configuration with the zero camber angles, and so with the wheels perpendicular to the ground. The bottom photograph on Fig. 3 presents the CENTAURO robot adopting the configuration with the non-zero camber angles, and so with the wheels at an angle with respect to the ground surface. Let us define the \u2018wheel plane\u2018 that is constructed upon the torus centre line, i.e. it divides the wheel in half with the plane normal being parallel to the wheel axis (yw \u2208 \u211c 3). Then the position vector from the wheel centre to the SPV (wxcp \u2208 \u211c 3) is, as shown in Fig. 2, wxcp = \u2212zcR \u2212 nr, (5) where R \u2208 \u211c + represents the torus major radius, r \u2208 \u211c + stands for a torus minor radius, n \u2208 \u211c 3 symbolises the unit vector along the ground normal, and zc \u2208 \u211c 3 denotes a unit vector that lies in the cross-section of a wheel plain and a plane constructed with the wheel axis and ground normal. The vector zc in (5) is not fixed in any of the robot frames. It is convenient to define a new frame at the wheel centre, referred as wheel-ground frame and marked by \u2018\u2018c\u2019\u2019 in this work (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000098_s10514-019-09887-8-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000098_s10514-019-09887-8-Figure3-1.png",
        "caption": "Fig. 3 First critical situation",
        "texts": [
            " In case of static obstacles, the process is similar except for the responsibility for avoiding collisions: the agent must take all the responsibility for avoiding collision. The same hold for dynamic agents that use different policy for planning their trajectories. In this section we will highlight a couple of scenarios where the original ORCA algorithm is not able to generate a set of collision-free trajectories leading each agent to its final destination. For both cases we will propose a solution. Figure 3a shows an example where agents A and B want to reach each other\u2019s starting position. Unfortunately the agents will never achieve their goals. Due to the alignment of the nominal velocities vA = vnomA and vBA = vnomBA the collision cone CC\u03c4 AB is like the one in Fig. 3b where the vectors \u03b7, u and the relative velocity (blue vector) are aligned with vA and vB . Since the vector \u03b7 gives the angular coefficient of the line that will create the half plane ORCAAB , we have that such a plane will be orthogonal to vA. The ORCA\u03c4 AB half-planes for agents A and B are shown in Fig. 3c. The new velocity vnewA will be selected by solving a linear programming prob- lem to obtain a velocity as close as possible to vnomA . Linear programming will choose vnewA as a vector that satisfied the constraints imposed by the ORCAAB planes and that minimizes the distance between vnewA and vnomA . Since the shortest distance between two points is given by the segment connecting them, and vA and vnomA are aligned, the result, i.e. vnewA , will still have the same direction and will only change the module. The two agentswillmaintain the samedirection ofmotion: when they get closer, the algorithm will gradually decrease their speeds till they stop, facing each other, remaining motionless. It is worth highlighting that if B is a static obsta- cle or a dynamic obstacle with constant velocity vB , the situation does not change much. Remark 1 It is worth highlighting that with noisy measurements the geometric conditions depicted in Fig. 3b should rarely happen. Nevertheless, we modify the algorithm in order to work also in noise-free and no-uncertainty ideal conditions. Another problem occurs when a group of obstacles ismoving in a close formation and another agent has tomove in the same line but with opposite direction. The problem arises because the ORCA method considers obstacles individually and so the superposition of half-planes may create deadlocks. Let A, B and C be agents moving according to the velocities in Fig. 4a. The agent A\u2019s goal, pgoalA , is behind two dynamic agents/obstacles, B and C , which are quite close to each other"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001207_rpj-07-2018-0171-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001207_rpj-07-2018-0171-Figure11-1.png",
        "caption": "Figure 11 Virtual comparison of milled component and the prototype",
        "texts": [
            " The deformation values within the tested loading range show linear behaviour. This means that deformationswere purely elastic. Because of the problem with the bolt, the axle carrier was modified once more (seventh design stage \u2013 optimization based on experimental data). The span between the steering rod mounting points was reduced, so the elastic deformation of the bolt was smaller. The model was evaluated only by FEM analysis, which showed the maximal deformation of 0.4mm. Virtual comparison of the milled component and the seventh design stage is shown in Figure 11. The design process was created and validated. Better results can be achieved if the topology optimization is carried out in several steps (three runs of topology optimization are relevant for this type of problem). Good reason for stopping the optimization is when the topology optimization starts to remove the material from areas, where it is necessary; The manufactured SLM prototype had satisfactory results with the geometrical deviations of less than 0.5 mm; For the cornering load case, results from photogrammetry measurement were similar to those obtained in FEM analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000336_j.ibiod.2015.04.023-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000336_j.ibiod.2015.04.023-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of the single air-cathode MFC configuration in this study.",
        "texts": [
            " The experiments to investigate application as a substrate and biodegradation of ETAwere conducted by comparing with acetate and conventional anaerobic treatments. To evaluate PP felt as a separator in a MFC using ETA, an experiment was performed with PEM, CEM, and PP felt. Using PP felt for ETA degradation in a single air-cathode MFC was first investigated, and the performance of the fuel cell was discussed. A rectangular acrylic reactor which has a single chamber consists of anode and cathode (a working volume of 75 mL following dimension: length, 3 cm; width and height, 5 cm) as Fig. 1. Carbon cloth (1071 HCB, AvCarb\u00ae) was used as both the anode (without wet-proofing) and the cathode (30% wet-proofing). One side of the cathode contained 0.5 mg cm 2 of Pt catalyst from 10% Platinum/ Carbon (10% Pt on Vulcan XC-72, Premetek) with Nafion solution as a binder. The air-facing surface of the electrode was coated with a carbon base and polytetrafluoroethylene (PTFE) in four diffusion layers. The cross-sectional area of the electrodes was 25 cm2, and they were connected by titanium wire"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001868_med48518.2020.9183031-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001868_med48518.2020.9183031-Figure1-1.png",
        "caption": "Fig. 1. The FRD (front, right, down) body-coordinate frame is shown on the left. The tilt angle \u03c7 represents how much the propellers are tilted, e.g. 0 deg corresponds to helicopter mode and 90 deg describes the airplane configuration. The individual thrust of each propeller is denoted by ti.",
        "texts": [
            " Most publications lack practical experiments to support their claims and the controller designs are mostly only validated in self-written simulations. The tiltquadrotor VTOL aircraft at hand differs from the common designs because it is both a quad-rotor and a tilt-rotor VTOL. A complete model has first been presented in [8]. This section describes the novel FPID controller designed for the tilt-quadrotor VTOL aircraft. First the notation and coordinate frames are introduced and then the controller architecture is presented in detail. In this paper, bold symbols indicate that the considered quantity is a vector. Fig. 1 shows a sketch of the VTOL aircraft and introduces the FRD (front, right, down) bodycoordinate frame denoted by a subscript \u2018B\u2019. The tilt angle of the propellers \u03c7 is measured w.r.t the body x-axis. The amount of thrust each propeller produces is denoted by ti and their sum, the total thrust, is denoted by T . If the actual direction of the thrust plays an important rule, the vector notation T is used and the following holds: T = TxB 0 TzB  = T \u00b7  sin(\u03c7) 0 \u2212 cos(\u03c7)  = 4\u2211 i=1 (ti)  sin(\u03c7) 0 \u2212 cos(\u03c7)  The negative sign for the thrust in zB-direction is because the thrust produced is pointing upwards, whereas the zBaxis points downwards"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001586_jsen.2020.2983261-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001586_jsen.2020.2983261-Figure3-1.png",
        "caption": "Fig. 3. (a) Image of monolithic TX with stacked CMOS drive and periodic loaded MZM. The MZM has a total push-pull RF/optic interaction length of 6 mm (3 mm per MZM arm). (b) Magnified view of periodically loaded electrode in MZM.",
        "texts": [
            " An MZM electro-optic modulator is used in the TX due to its relatively temperature insensitive operation. The MZM has lateral PN junctions with 4 \u00d7 1017 cm\u22123 peak p and n doping concentrations (as indicated by simulations of the processing implant conditions and thermal budget), a \u223c135 nm SOI thickness, and 2 \u03bcm BOX appropriate for a \u223c1.3 \u03bcm wavelength of operation. The pn junctions show a leakage current on the order of nanoamps with a 1 V reverse bias, and so have a shunt resistance on the order of hundreds of mega-ohms. An image of the TX is shown in Fig. 3(a). The CMOS driver differential output is coupled directly into a periodically loaded push-pull MZM. The MZM incorporates unloaded RF electrode segments to increase the electrode effective line impedance. The unloaded RF segments are transmission line sections that are not connected to pn junctions and are periodically inserted into the MZM electrode. Fig. 3(b) shows the loaded sections (electrode sections coupled to a pn-junction) are the straight horizontal electrode segments highlighted by thick white arrows, and the unloaded sections are the 180\u00b0 turn segments denoted by the narrow white arrows. Each MZM arm has ten 300 \u03bcm long electrode sections loaded with optical pn-junctions (3 mm total loaded electrode length in each MZM arm) and nine 157 \u03bcm long unloaded elec- trode sections, resulting in a \u223c68% loading and 4.41 mm total electrode length. Since the unloaded electrode sections have low capacitance their relative contributions to RF propagation losses are small"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000492_s12206-016-0115-8-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000492_s12206-016-0115-8-Figure1-1.png",
        "caption": "Fig. 1. Planetary gear system coordination.",
        "texts": [
            " d r , \u03b8, F r r and M r are vector variables representing the 3-dimensional translation, rotation, forces, and moments of the stations, respectively. The subscripts represent the component (a) or connection relationship (b). The sun gear, planet gear, ring gear, and carrier are represented as \u2018s\u2019, \u2018p\u2019, \u2018r\u2019 and \u2018c\u2019, respectively, and the relationships between the sun gear and the planet gear, between the planet gear and the ring gear, and between the carrier and the planet gear are represented by \u2018sp\u2019, \u2018pr\u2019 and \u2018cp\u2019, respectively. The superscripts represent the station number (n). Fig. 1 illustrates the motion relationships of a planetary gear system. A global fixed-coordinate system C is set, and each part has its own coordinate system. s pi rC , C , C and cC are the local fixed coordinates of the sun gear, i-th planet gear, ring gear, and carrier, respectively. In the ideal condition, each coordinate has a specific relationship s pi r(C C C C= = = = cC ) and shares a unit vector ( i , j, k r r r ). Each variable can be represented with a harmonic equation, as in Eqs. (1) and (2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000158_j.mechmachtheory.2019.103625-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000158_j.mechmachtheory.2019.103625-Figure2-1.png",
        "caption": "Fig. 2. A single limb showing (a) the reference and (b) the j th pose.",
        "texts": [
            " . , m . All displacements are given with respect to the reference pose given by r 0 = 0 and \u03c60 = 0 . At the reference pose, the rotation angle \u03b8 j = \u03b80 = 0 . It is noted that mechanism in this synthesis is obtained by constraining 3-RRR planar parallel linkage. By this way, the number of degree of freedom is reduced from 3 to 1. Moreover, the obtained 10-bar planar mechanism has a smaller design space than the ordinary 10-bar linkages [25] , due to the presence of constraining parallelograms. Fig. 2 displays a single limb of the 3-RRR planar parallel mechanism. Without loss of generality, the planar parallel mechanism that is used for synthesis is of arbitrary configuration, not necessarily symmetric. For dyad A 1 C 1 , by virtue of the link rigidity, the length remains constant during motion. This means (a 1 j \u2212 c 1 j ) T ( a 1 j \u2212 c 1 j ) = l 2 1 , j = 1 , . . . , m (1) where a 1 j and c 1 j are position vectors of joints A 1 and C 1 at the j th pose, l 1 is the link length of A 1 C 1 . The equation is applicable to other two dyads, A 2 C 2 and A 3 C 3 , of Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001460_s00170-019-04874-w-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001460_s00170-019-04874-w-Figure9-1.png",
        "caption": "Fig. 9 Experimental setup methods for powder flow image",
        "texts": [
            " The minimum volume is 0.02 (mm3), and the maximum volume is 4.13 (mm3). Meanwhile, the quality coefficient of all elements are above 0.4, indicating good quality for fluid analysis. Finally, the calculation is performed to obtain a time-averaged stationary solution. An experiment of photographic system is built to measure the concentration distribution and verify the powder flow process compared with the calculated analysis. The experimental setup methods for powder flow image are illustrated in Fig. 9. The parallel laser sheet with width 0.8 mm is generated by a 532-nm green pulse laser. Images are captured in a dark background with a CCD camera at 30 frames per second to improve resolution. The particle cloud injected from the nozzle absorbs the light intensity of the laser sheet, and then the camera can capture the reflected image of the particles on the laser sheet plane. The post-image processing method is based on the Mie theory, which describes the relationship between the particle image gray level and the actual corresponding particle concentration [23]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001759_j.mechmachtheory.2020.103995-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001759_j.mechmachtheory.2020.103995-Figure5-1.png",
        "caption": "Fig. 5. Double spherical 6 R linkage acts as a crank-rocker. (a) Configuration of limit position at point M and (b) configuration of limit position at point N.",
        "texts": [
            " 4 shows the variations of \u03b86 as functions of the input angles \u03b81 . The results also show that this kinematic path has two distinctive motion modes under different design parameters. The linkage converts the uniformly rotating motion of the input link 12 into the oscillating motion of the output link 56 under the conditions of | a ( r 2 \u2212r 1 ) tan \u03b1 | < 1 , i.e., it is a crank-rocker linkage. However, both two rotating links make complete revolutions, under the conditions of | a ( r 2 \u2212r 1 ) tan \u03b1 | \u2265 1 , i.e., double crank linkage. Fig. 5 describes two limit positions of crank-rocker. In both cases, the axes of four revolute joints 1, 2, 4 and 5 are coplanar. The configurations of the two limit positions, also called dead center positions in crank-rocker linkage, correspond to points M and N in the kinematic paths of Fig. 4 respectively. The maximum swing range \u03b86 max of the rocker can be expressed as \u03b86 max = 4 \u2223\u2223\u2223\u2223arctan ( a cos \u03b81 lim ( r 2 \u2212 r 1 ) sin \u03b1 \u2212 a cos \u03b1 sin \u03b81 lim ) \u2223\u2223\u2223\u2223. (27) Unlike the maximum swing range of planar or spherical crank-rocker is less than \u03c0 , this spatial crank-rocker can have oscillation angles greater than \u03c0 , i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000477_0954406216631781-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000477_0954406216631781-Figure1-1.png",
        "caption": "Figure 1. The structure of the motorized spindle.",
        "texts": [
            " The simplified heat transfer coefficients of the spindle surface can be given by19 ht \u00bc 28 1\u00fe ffiffiffiffiffiffiffiffiffiffiffiffi 0:45vt p \u00f022\u00de where vt is the velocity of the spindle surface, as for the end region of the rotor and the spindle, vt is the average velocity of their end region. The heat transfer of the motorized spindle surface includes convection and radiation. The entire thermal resistance can be seen as parallel two thermal resistances, so the entire heat transfer equals the sum of the convection and radiation heat transfer \u00bc c \u00fe R \u00f023\u00de where is the entire heat transfer, c is the convection heat transfer, R is the radiation heat transfer. According to Chen and Chen,20 the heat transfer coefficient of the motorized spindle surface is 31.6W/(m2 C). Figure 1 shows an air-cooled motorized spindle, whose power is 8KW, while the operating voltage is 380V, and the rotating speed is 6000 r/min. The spindle is supported by the front and the rear bearings. The front bearings are two angular contact ball bearings (7010C/P4) and adopt the face-to-face mounting. The rear bearing is a deep-groove ball bearing (6007) and is compressed by a wave spring. The wave spring can compensate the thermal deformation of the spindle. The lubricating grease is Kluber L252. The spindle cooling fan is a stamped product with six fan blades mounted on the spindle by four hexagonal head bolts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001688_s11661-020-05831-z-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001688_s11661-020-05831-z-Figure3-1.png",
        "caption": "Fig. 3\u2014Computational domain used to analyze line energy density.",
        "texts": [
            " Table II shows the required physical properties of 316L stainless steel.[42] In addition, the laser spot diameter was 54 lm. B. Influence of Line Energy Density on the SLM Formation Process To comprehensively consider the influence of laser power and scanning speed on the SLM formation process, the line energy density Q was analyzed where Q \u00bc Plaser=v. To facilitate the comparison of the effect of line energy density on the SLM process, the calculation used here was the SLM single-layer single-pass formation process. Figure 3 shows the calculation domain used to analyze the line energy density. The overall size was 1000 lm 9 150 lm 9 130 lm, where the substrate thickness was 50 lm, and the mesh size was 2.53 lm3. The powder bed distribution in Figure 3 was calculated by Yade wherein the average particle radius was 25 lm, and the half-width of the particle radius (the difference between the average and the maximum or minimum) was 5 lm. The powder bed had a thickness of [16] METALLURGICAL AND MATERIALS TRANSACTIONS A 50 lm, and the obtained powder bed had a packing density of 58.47 pct. During this calculation, the initial temperature was 300 K, and the laser started moving from the position of the horizontal coordinate (100, 75 lm). It stopped when it reached the defined position (900, 75 lm). The domain was then cooled for 100 ls. In terms of computational efficiency, the average calculation time required for the SLM single-pass example was 8 hour. Figure 4 shows the simulation results of temperature field and solid fraction distribution at different times when the line energy density was 200 J/m; here, the sectional view is the Y-direction middle section of the calculation domain shown in Figure 3. The side thermal boundary condition was symmetric and had less influence on the temperature field of the molten pool. The metal particles were melted by heat to form a molten pool once the laser began to act on the powder bed; as the laser moved forward, new particles were continuously filled in the front of the molten pool, and the tail of the molten pool gradually cooled down to form a solidified track. A part of the substrate was also melted by heat and solidified. Finally, the laser action area formed a continuous solidified track in the powder bed",
            " Therefore, a moderate line energy density should be used to obtain a relatively flat solidified track and establish a good connection with the substrate or the former layer. C. Influence of Hatch Space on the SLM Formation Process To analyze the effect of hatch space on the SLM formation process, the calculation used here was the SLM single-layer multi-pass process. Figure 11 shows the calculation domain used to analyze the hatch space. The overall size was 1000 lm 9 400 lm 9 110 lm where the substrate thickness was 30 lm, the powder bed thickness was 50 lm, the mesh size was 2.5 lm, and the particle size distribution was consistent with that in Figure 3. During the calculations, the laser was applied in five paths in sequence. During the process of forming each path, the laser moved from the X-coordinate 100 lm to X-coordinate 900 lm, and the Y-coordinate difference of adjacent paths was the hatch space. The laser power and scanning speed were maintained at 240 W and 1.5 m/s. We note that the calculation domain size used herein was limited, and the length of laser single-pass formation in the actual SLM process is often several tens of millimeters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003626_piae_proc_1922_017_035_02-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003626_piae_proc_1922_017_035_02-Figure4-1.png",
        "caption": "FIG. 4.",
        "texts": [
            " (18a) If the car is accele- rated so that p is much greater than n we have a deflection of the spring equal to, or less than, the height of the hump. This assumm that there is no jumping of the wheel, which is hardy allowable at high speeds. On the other hand, we w e from equation (18) that slowing down will reduce the added spring stress very rapidly untii at rest it beoomes, of course, zero. PART V .-DOEBLY-SPRUNG WAGON-V~CTORIAN BRoUr;H.ur 4r;n THE RAILWAY WAGON. Assume now that we have two axles each connected sepamtdy This separate by a spring or springs with the body, see Fig. 4. at The University of Auckland Library on June 5, 2016pau.sagepub.comDownloaded from PRINCIPLES OF VEHICLE SUSPENSION. 47 1 springing of vehicles was the epoch-marking invention of Obadiah Elliot (1801), by which unsprung weight was reduced and permissible or safe speeds increased. Previously the perch-pole of great mass inherited from the so-called Ancient Britons tied the axlw together and formed the back-bone of the four-posted frame froni which the body was slung by leathern straps carried over quarter-elliptic springs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000477_0954406216631781-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000477_0954406216631781-Figure7-1.png",
        "caption": "Figure 7. Schematics of the sampling points: (a) surface development figure; (b) back view of the motorized spindle; (c) sketch of the sampling planes.",
        "texts": [
            " The rotating speed and the vibration are measured by the overall dynamic balance instrument (wiBalancer). The noise is measured by the digital sound level meter (CEM DT-8852). Figure 6 shows the experimental setup. The ambient temperature is 27 C and the temperature of the entire motorized spindle surface is assumed to be the same. The natural air flow in the lab is less than 0.1m/s. In order to make sure the rotating speed exactly meets the test requirements, the speed sensor is set at the front of the spindle to measure the rotating speed. Figure 7(a) shows the map of the surface sampling points. There are nine lines of the sampling points and each line has 30 sampling points, with the 270 sampling points in total. As shown in the Figure 7(c), there are nine sampling point lines along the Y direction dividing the motorized spindle into nine sampling planes, and each plane is divided evenly by the 30 points (30 sampling points along the X direction). Among all the points, there are five key sampling points (points 1 (15,1), 2 (15,5), 3 (15,9), 4 (25,5), and point 5 (5,5)). In the temperature rise experiments, the temperature of the key points is recorded each minute. In the experiment of the steady-state temperature, the temperature of every sampling point is also recorded"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003678_a:1008896010368-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003678_a:1008896010368-Figure2-1.png",
        "caption": "Figure 2. Leg actuation.",
        "texts": [
            " Each foot is in contact with the ground along its entire lower surface. A2. The pitch angle of the trunk, q5, is fixed during walking. Assumption A1 implies that it is possible to supply an arbitrary input torque to the robot through the ankle joint. Assumption A2 means that the inertia force due to the rotation of the trunk is negligible in the dynamic equations of the robot. These assumptions are realized by constructing each leg using a pair of parallel links. The structure and actuation of the leg are shown in Fig. 2. Two D.C. motors attached at the trunk supply input torques \u03c41a and \u03c42a in Fig. 2(a) through a timing belt and reduction gear. Since \u03c41a = \u03c41b and \u03c42a = \u03c42b, the leg can be modeled simply as shown in Fig. 2(b). The constraints are expressed as c(q) = 0. (1) Since the right foot in Fig. 1 is fixed on the ground, c(q) = [ l1 sin(q1)+ l2 sin(q2)\u2212 l2 sin(q3)\u2212 l1 sin(q4)\u2212 L l1 cos(q1)+ l2 cos(q2)+ l2 cos(q3)+ l1 cos(q4) ] . (2) The dynamic equation for the robot is written as H(q)q\u0308 + h(q, q\u0307)+ CT (q)\u03bb = \u03c4, q = (q1, q2, q3, q4) T , H \u2208 R4\u00d74, h \u2208 R4\u00d71, \u03bb \u2208 R2\u00d71, \u03c4 \u2208 R4\u00d71, C \u2208 R2\u00d74, (3) P1: VTL/TKL P2: EHE/TKL P3: VTL/TKL QC: PMR/TKJ T1: PMR Autonomous Robots KL465-06-Mitobe May 16, 1997 17:19 Control of a Biped Walking Robot 289 where Hq\u0308 represents the inertial forces, h is the gravitational, centripetal and Coriolis term, \u03bb represents the constraint forces, CT\u03bb represents the torques at each joint due to the constraints, and \u03c4 represents the joint torques"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure3-1.png",
        "caption": "Fig. 3. speed x",
        "texts": [
            " Generally, the location of the centre of mass is unknown; but, even assuming that another test is conducted to determine the centre of mass location, it can still be impractical to assume that the body can be positioned with sufficient precision so as not to introduce significant error in the inertia estimate by misalignment [18]. As well, the multifilar pendulummethod may suffer from decreased accuracy due to the linearization required to obtain Eq. (6). The multifilar pendulum can also experience side swaying that reduces estimation accuracy [18]. The compound pendulum can be used when it is difficult to align the centre of mass with the axis of rotation. The object is hung by wires rotating on two overhead points (Fig. 3a). The moment of inertia about an axis parallel to the two overhead points and through the centre of mass can be calculated by ICP \u00bc ml2c gT2 4p2lc 1 ! \u00f07\u00de (a) A compound pendulum oscillates an arbitrary object hung by wires rotating on overhead pins. (b) An arbitrary object spins at a constant angular C about an axis through its centre of mass. where T is the oscillation period, g is gravity, m is the mass of the object, and lc is the distance between the axis of rotation and the line between the two points of support [16]",
            " Typically, for the compound pendulum, the distance to the centre of mass from the rotation axis cannot be easily determined, especially when the centre of mass point is inaccessible and so a direct measurement is not practical [16]. Eqs. (5)\u2013(7), are very sensitive to the measurements of T and other parameters (l; L; r, and ks). For high accuracy inertia estimates, T needs to be measured with uncertainties on the order of less than a millisecond. A method to estimate the products of inertia is by spinning the object at a constant angular speed about an axis through its centre of mass [16], as shown in Fig. 3b. From Eq. (4), if the object is spinning with constant angular speedxC \u00bc _h t\u00f0 \u00de such that _xC t\u00f0 \u00de \u00bc 0, the products of inertia can be determined from the reaction torques, sx and sy, and xC as IOyz \u00bc sx x2 C \u00f08\u00de IOxz \u00bc sy x2 C \u00f09\u00de For small products of inertia, high angular speeds may be necessary to measure reaction torques accurately. Practically, there may be constraints on the maximum speed that the object can withstand; and measurements of reaction forces at the bearings supporting the object requires additional instrumentation such as force transducers mounted perpendicularly to each other"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000447_jmes_jour_1960_002_027_02-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000447_jmes_jour_1960_002_027_02-Figure1-1.png",
        "caption": "Fig. 1. Tube ironing using one dividing plane",
        "texts": [],
        "surrounding_texts": [
            "MUCH RECENT PROGRESS has been made in the derivation of upper-bound solutions for problems concerning the plane strain deformation of rigid, perfectly plastic materials. Hill (I), Green (2), and Johnson (3)$ have developed a method using discontinuous velocity fields based on the maximum work inequality derived by Hill (I). This method is limited in its application since only perfectly smooth and perfectly rough boundaries can be discussed at present. In the majority of practical cases, some portion of the boundary is acted upon by Coulomb friction and the method described below has been developed to give solutions for certain problems, taking Coulomb friction into account. Upper-bound solutions for tube ironing, sheet drawing and extrusion, and plane compression are considered. The first example, tube ironing, is studied in detail to explain the origin of the force field method. The material considered is rigid, perfectly plastic. For stresses below the yield point, the elastic strains are zero, and for stresses which cause yielding the material flows without strain hardening. Some allowance must therefore be made if the proposed solutions are applied to strain hardening materials. Notation F\u2019 H \u2019 h The M S . of this paper was received at the Institution on 1st April 1960. * New Conveyor Company Limited, Smethwick, Birmingham; formerly holder of Tube Investments Research Scholarship, Department of Mechanical Engineering, University of Shefield. Graduate of the Institution. t Senior Research Fellow, Department of Mechanical Engineering, University of Shefield. $ A numerical list of references is given in Appendix II. External forces per unit width. Initial thickness of tube or half-thickness of sheet Final thickness of tube or half-thickness of sheet. or lamina. J O U R N A L M E C H A N I C A L E N G I N E E R I N G S C I E N C E Yield shear stress in plane strain. Normal force per unit width. Operating force per unit width. Tangential force per unit width. Half-width of lamina in plane compression. Dimensionless factors. Die angle. Coefficient of Coulomb friction. Normal stress. Shear stress. Yield stress in uniaxial tension. Variable angles in the physical configuration."
        ]
    },
    {
        "image_filename": "designv11_14_0003626_piae_proc_1922_017_035_02-Figure17-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003626_piae_proc_1922_017_035_02-Figure17-1.png",
        "caption": "FIG. 17. F I G . 18.",
        "texts": [
            " This has been very well known for years, and it became a scientifio fact as a result of Rairstow's work an the elastic limit of steel and Dalby's indieator work on internal friction of strains. FOP a solid steel spring either a single plate, helix or torsional tube, the load-defleotion curve is a straight line, and any mass attached to the spring vibrates with simple harmonic motion of constant ainplitudo. If a solid steel spring be coupled to a dash-pot so that them is a fluid frictional iwsistanae proportional to speed of deflection, the load-deflection diagram for slow iiiotion or static tests is a straight line, and for quick iiiaintaiiied motions we have looped curves as in Fig. 17 . Any mass attached to the spring vibrates in smooth harmonic osallation., of which the amplitudes decrease in geometrical progression; 3ee dppendix n. I n all cases they are elliptic curves. l f now a spring of any sort be coupled to a solid friatiion device, a dry guide on a dry slide-bar, say, the lod-esten,Gon diagram is altmed in charaotRr and is of the form shown in Fig. 18. Any inas< attache(d to such a spring vibrates in harmonic motions, of which the amplitudes decrease in arithmetical progremios and of which the mid points thane from om to the other Bide of ithe zero-see Appendix E. The conditioas of Fig. 18 a m realised in a light-car suspension with dry springs excessively nipped, and the aonditions of Fig. 17 aic realised in oertain damped governors of the old Pamoins type and ir, an approximate way in the ,well-lubrioated springs of high-grade oars when newly assembled. There are one or two important variatio,ns on these main types. Conbider, for iiistanae, locomotive springs, which usually rim dry aid have very little nip. The friction b'etween the plates is of the solid type and varies with the load. The diagram is of the at The University of Auckland Library on June 5, 2016pau.sagepub.comDownloaded from 502 THE INS'I'I'I'U1'ION O Y A U1'O~IOHILE ENGINEERS form shown in Fig",
            " A further variation, and the last that we need oonsider before prooeding to generalisation, is the so-called shock-absorber, which as a rule memly adds fluid friction to the system. This is characterised by adding the diagrams in Figs. 17 and 18, see Fig 22. and it is obvious at a glancle that shook-absorbers of this type aro useless on springs which already h a w much solid friction. Shock-absorbers of the fluid-friction type are useful as remedies OIL mrs where the springs have iiisufficient pip or few leaves, i4 which case they inay approach Fig. 17 They are useful in special circumstances, as in racing cays, but they are quite unnecessaiy ab standttrd fittings to Standard touring carb ii care is used in designing the springs. Here it may be added that almost the same kind of opinion holds regarding auxiliary helical springs. These a x useful when nip is exclessive o r where leaves are many or narrow where the main springs are unduly stiff. They am palliatives I n Mr. Baillie\u2019s interesting paper on \u2018\u2018 Springs,\u201d\u201d a number of facts are recorded R hich are well worthy of thought"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001978_0954407020964625-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001978_0954407020964625-Figure5-1.png",
        "caption": "Figure 5. Planetary gearbox fault implantation bench.",
        "texts": [
            " The testing process as follows: Each test sample are averagely segmented at first, and then the trained L1/2SF model is adopted to calculate the local features from the segments. Next, the feature vector of each test sample are also obtained by the average way. At last, the learned feature vectors are input to the trained softmax classifier for fault identification. Fault diagnosis cases under variable rotational speed Case 1: fault diagnosis of a planetary gearbox Data description. A planetary gearbox fault implantation bench is adopted for signal acquisition under variable rotational speed. As exhibited in Figure 5, the test bench contains a motor, a planetary gearbox, two bearing seats and two shaft couplings. Three planet wheels are designed in the gearbox. The teeth numbers of the sun wheel and planet wheels are 36 and 18, and the module n is 1.5, respectively. There are ten health conditions as shown in Figure 6: normal condition (NC); three sun wheel fault types (crack, pit and worn tooth), which are named as WC, WP and WW; three pinion fault types (crack, pit, and worn tooth), which are named as PC, PP and PW; and three coupled fault types (wheel worn and pinion worn, wheel pit and pinion crack, wheel pit and pinion worn), which are named as WWPW, WPPC and WPPW"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.19-1.png",
        "caption": "Fig. 10.19 The structure of centrifugal acid removal equipment",
        "texts": [
            " The waste acid after the nitration can be naturally filtered through the NC layer, by which the waste acid recovery rate is only about 60%; acid removal by centrifuge can drive out a large number of concentrated waste acid in NC so that the recovery rate of waste acid reaches more than 90%. After removing acid, NC should be quickly immersed in water to make an NC solution with a concentration of 2% (2 g/100 ml H2O) or below, which facilitates the pumping for the digestion process. (1) Upper suspension centrifugal acid removal equipment Centrifugal acid removal equipment has a structure of upper suspension and bottom discharge, of which centrifuge is made of acid-resistant stainless steel. Figure 10.19 shows the structure of the upper suspension and the lower discharge centrifugal acid removal equipment. Before feeding of NC, the front cover and cone cover 5 of centrifugal acid removal equipment are closed. The rotation rate of the centrifuge is adjusted to 200\u2013350 r/min while feeding; after loading, the centrifuge is set as 700\u2013 1000 r/min to drive acid. After a certain time, stop the centrifuge and lift the cover. The NC on the screen is divided into small pieces manually by aluminum fork or shovel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001354_icems.2019.8922412-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001354_icems.2019.8922412-Figure1-1.png",
        "caption": "Fig. 1. Coordinates of PMSMs",
        "texts": [
            " This work was supported by the basic operating expenses of central scientific research (PA2018GDQT0021). 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE II. PMSM MODELING A. The Mathematical Model of PMSM In this section, a generalized voltage model for both Interior Permanent Magnet (IPM) Synchronous Motors and Surface Permanent Magnet (SPM) synchronous motors is introduced, where proper significance of the PM flux linkage is assumed, that is used to design a linear Luenberger observer. As is shown in Fig. 1, where ud,q , id,q , Ld,q are the components of voltage, current and inductance of the equivalent dq-circuits , u\u03b1,\u03b2 , i\u03b1,\u03b2 , L\u03b1,\u03b2 are the components of voltage, current and inductance of the equivalent \u03b1\u03b2-circuits, \u03c9e is the electrical speed of PMSM,\u03b8e is the rotor pole position electrical angle. The model of PMSM can be presented in the d-, qreference frame as follow: 0 = + s d e q d d q q e f e d s q d R L Lu idt u id L R L dt \u03c9 \u03c9 \u03c8 \u03c9  + \u2212                    +    (1) where Rs is the phase resistance, f\u03c8 is the rotor PM flux linkage"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001184_j.ijfatigue.2019.105281-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001184_j.ijfatigue.2019.105281-Figure5-1.png",
        "caption": "Fig. 5. Setup for the mode II crack growth test.",
        "texts": [
            " 4a) and side grooves (detail B shown in Fig. 4a). The fatigue crack is initiated at the tip of the chevron notch. The 60\u00b0 V-shaped side groove is designed to cause mode II fatigue crack growth along the neutral section. For a constant amplitude load, \u0394KII of the crack along the neutral section gradually decreases with the growth of fatigue crack. The mode II threshold stress intensity factor range \u0394KII,th can be calculated if the length of the arrested crack is measured in the test. The setup for the test is shown in Fig. 5. A pair of specimens was tested at the same time. The upper and lower cantilevers of each specimen were clamped between two grips using two M12 bolts. A load cell was placed between the grip and bolt head to measure the clamping force, which was set to 10.58 kN in the test. Two cylinders with diameters of 2mm were inserted between the specimen and grips, ensuring that the line loading is applied to the specimen during the test. A third cylinder with a diameter of 1mm was inserted in the gap between the two cantilevers. It was used to apply an identical force to the upper and lower cantilevers. The mode II crack growth test was performed using a tension\u2013compression fatigue test machine. The specimens were held by the setup shown in Fig. 5 and then connected to the testing machine. The tests were performed at a constant load amplitude of\u00b1 5.9 kN and at a frequency of 6 Hz, with a sinusoidal waveform. Trial tests show that the mode II fatigue crack stops to grow after 2\u00d7106 loading cycles. Therefore, the test was stopped when the number of loading cycles reached 2\u00d7106. After the test, the specimen was sectioned along the centre. The crack growth path on the polished section was observed using an optical microscope. Fig. 6 shows the scrapped CL60 railway wheel, which was subjected to subsurface RCF"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002993_j.msea.2021.141494-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002993_j.msea.2021.141494-Figure8-1.png",
        "caption": "Fig. 8. Average hardness and associated error (expressed as the standard deviation of measured dataset) calculated from the macro-, micro- and nanoindentation techniques from the three surfaces (front, side and top) of the LENSTM-processed specimen.",
        "texts": [
            " However, instead of a gradual reduction along the building direction, variation in hardness is sporadic for these two surfaces, therefore causing a large error (expressed as the standard deviation) in the average value, especially for the side surface. Such sporadic variation in hardness with large standard deviation in the average value appears for the top surface as well. Furthermore, the hardness measured by macro-, micro- and nanoindentation techniques (at normal loads of 1471 N, 4.9 N and 5 mN, respectively) from the three surfaces (front, side and top) of the DEDprocessed specimen are compared in Fig. 8 in order to better identify the effect of local microstructural inhomogeneity on the mechanical response. The macro- and micro-hardness are nearly equal for the three surfaces, whereas both average hardness and standard deviation in nano-indentation are much higher. The macro-hardness from front, side, and top surface are 4.27 \u00b1 0.11 GPa, 4.53 \u00b1 0.17 GPa, and 4.01 \u00b1 0.18 GPa, respectively, whereas nano-hardness for these surfaces are 9.66 \u00b1 3.52 GPa, 9.12 \u00b1 3.15 GPa, and 9.92 \u00b1 4.35 GPa, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001152_ab3f56-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001152_ab3f56-Figure2-1.png",
        "caption": "Figure 2. Configuration of outer and inner sleeve shell materials. Production drawings of inner sleeve (A), and outer sleeve (B). (a) Elastic band, (b) Spun-silk fabric, (c) Single-layer mesh fabric, (d) Nylon band.",
        "texts": [
            " Elastic bands are used for padding on both ends of the inner shell for stable attachment of the device to the targeted body part during wearing. The outer layer generates the compression force to the body by pulling the ends of the sleeve together by wires. As the outer shell needs to have proper stiffness in the direction of wire pulling, a single-layer mesh fabric was used. Two bars made of acrylonitrile butadiene styrene (ABS) were attached at the two ends of the outer layer to evenly distribute the pulling force from the wires. The detailed configurations and dimensions of the outer and inner sleeve shells are shown in figure 2. In order to measure the pressure applied to the skin by the device, a soft pressure sensor was integrated to the device. The soft pressure sensor was made of fabric to provide comfort for the user [16]. The sensor measures capacitance between two electrode layers which were made of knit jersey conductive fabric (Adafruit, NY, USA). The dielectric layer was fabricated using Ecoflex 30 (Smooth-Onm PA, USA). When the pressure is applied to the sensor, the distance between the two electrodes decreases and the capacitance increases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002671_j.jsv.2021.115967-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002671_j.jsv.2021.115967-Figure13-1.png",
        "caption": "Fig. 13. The corrected mode shapes estimated from the accelerometers, Eq. (26) and Table 2 : (a) original estimated mode shapes, (b) corrected mode shapes adjusted for tilt error.",
        "texts": [
            " 11 is mainly comprised of the first mode. The amplitude difference - based on the analytic expression for tilt - corresponds well with the ratio of standard deviation between accelerometers and lasers. We apply these amplitude differences to adjust the measured acceleration from the accelerometer, see Fig. 12 . This corrected the amplitude of the acceleration from the accelerometers so they have similar amplitude as the lasers. Similarly, we adjust the estimated mode shapes for the tilt effect in Fig. 13 where we obtain a better resemblance to the mode shapes from the model, see Fig. 9 . Here the MAC-value increases to above 0.99 for the first mode shape as compared to the values in Table 1 . Finally, we estimate the angular displacement by the method from from section 4.1 using the four vertical sensors (unaffected by translational motion) and we conclude that the assumption of small angles is valid for the given case, see Fig. 14 . This estimation of angular displacement only requires the vertical displacement and the geometrical distances to the centre of rotation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001662_j.addma.2020.101294-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001662_j.addma.2020.101294-Figure6-1.png",
        "caption": "Figure 6 Vertical deposition area as a function of width and height of the deposition",
        "texts": [
            " Material transfer by depositionMaterial loss due to solidification Mass change rate ( ( )) ( ) ( ) d V t A t v t f dt      (12) In the equation above, \u03c1, V, A and v are density, deposition volume, area and scan speed, respectively. The material transfer parameters are \u03bc representing the powder catchment efficiency and f representing the deposition rate. The deposition volume represents the whole volume of the melt pool and it is calculated based on the width, height, and length of the melt pool. The area represents the cross-section of the melt pool and it is calculated using the melt pool width and height, which are shown in Figure 6. In volume and area calculations below, the wetting angle (contact) of the melt pool is assumed to be 90o. ( ) ( ) ( ) ( ) 6 ( ) ( ) ( ) 4 V t w t h t l t A t w t h t     (13) In the equation above, w, h and l are the width, height and length of the deposited material, respectively. In this work, the material transfer rate \u03bcf is parametrized as a function of the laser power and critical laser power, instead of estimating the powder catchment efficiency [34]. The reason for the estimation approach is that the laser power is the dominant parameter, which defines the capacity how much powder is transferred to the melt pool"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001905_j.mechmachtheory.2020.104095-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001905_j.mechmachtheory.2020.104095-Figure15-1.png",
        "caption": "Fig. 15. Vibrating around axis x p .",
        "texts": [
            " In case I, the drive signals are set as six groups: 50V-5Hz(the amplitude of drive signal is 50V, the frequency of drive signal is 5Hz, the rest is similar), 50V-10Hz, 50V-30Hz, 100V-10Hz, 100V-30Hz, 150V-5Hz. Experimental results of case I are shown in Fig. 14 . The amplitude difference between numerical results and experimental results is shown in Table 4 . Errors and relative errors of case I are shown in Table 5 . Computational time comparison between the full model and the reduced-order model is shown in Table 6 . The simulation time is set as 0.5s. Case II: vibrating around axis x p . Fig. 15 shows the vibration mode of case II. The red chain dash-dotted arrow means the vibration direction of P 0 . \u201cSolidDotted\u201d arrow means the vibration direction of P 2 and P 3 . When P 2 vibrates along the direction of dotted arrow, P 3 also vibrate along the direction of dotted arrow. When P 2 vibrates along the direction of solid arrow, P 3 also vibrate along the direction of solid arrow. In case II, the sinusoidal drive signals are applied to the piezoelectric ceramic plates of kinematic chain II and kinematic chain III"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001138_1.g003926-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001138_1.g003926-Figure4-1.png",
        "caption": "Fig. 4 Pyramidal arrangement of four CMGs.",
        "texts": [
            " [36], the updated controller coefficient matrix\u0398k 1 that minimizes Eq. (29) is given by recursive least squares. Let P0 R\u22121 \u0398 and, for all k \u2265 1, let \u0398k 1 be the unique global minimizer of the retrospective cost function [Eq. (29)]. Then, \u0398k 1 is given by \u0398k 1 \u0398k \u2212 Pk\u03a6T f;k\u0393\u22121 k \u03a6f;k\u0398k Rz Ru \u22121Rz zk \u2212 uf;k (30) Pk 1 Pk \u2212 Pk\u03a6T f;k\u0393\u22121 k \u03a6f;kPk (31) where \u03a6f;k \u225c Gf q \u03a6k, uf;k \u225c Gf q uk, and \u0393k \u225c Rz Ru \u22121 \u03a6f;kPk\u03a6T f;k. V. Construction of Gf Consider a spacecraft actuated by n 4 CMGs mounted in the pyramidal arrangement shown in Fig. 4. Because the CMGs apply a pure torque to the bus, the location of the pyramidal arrangement on the bus is immaterial. The pyramid has a square base, for which the sides define the axes i\u0302B and j\u0302B of the bus. Consequently, the sides of the pyramidal base are aligned with i\u0302B and j\u0302B, and \u03b2 is the angle between a vector orthogonal to each face and the apex direction k\u0302B. The CMGs are arranged such that, for i 1; 2; 3; 4, each axis j\u0302Gi of rotation of gimbal Gi relative to the bus is perpendicular to the corresponding face; each axis i\u0302Gi of rotation of wheelWi relative to the gimbal is parallel with the corresponding edge of the base"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000994_tmech.2019.2916990-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000994_tmech.2019.2916990-Figure2-1.png",
        "caption": "Fig. 2. WFT system configuration and the mechatronics assembly [16].",
        "texts": [
            " This paper is organized as follows. Section II describes the principle of WFT. Section III elaborates the online calibration approach by modeling of a two-axis WFT. In Section IV, the proposed calibration is tested and compared with other methods. The system configuration of the adopted WFT is reviewed and introduced first, followed by the general principle of the WFT signal decoupling processes, from which the error models for WFT calibration can be derived. The adopted WFT is a block-designed system. As shown in Fig. 2, the system configuration includes the sampling module, the transfer module, and the host computer. In the sampling module, a large range MFS is self-built and integrated into the vehicle wheel. Detailed design and optimization techniques can refer to our previous cases studies [29], [30]. It has the rated loads about 50 kN force and 7.5 kN\u00b7m torque within an estimated accuracy error less than 2% and an observed dynamical frequency of the sensor about 300 Hz. An absolute photoelectric encoder is included to detect the rotation angle of the wheel. It has a speed limit about 2.1 \u00d7 103rad/s with a resolution of 0.08\u00b0 and 400 Hz signal output frequency. For signal processing (see Fig. 2), the force-voltage data are filtered and amplified by the conditioning circuit, then together with the angle signal, both of them are sent to the sampling channel of microcontroller unit (MCU). The MCU packs these data and sends them to the transfer module by wireless means. Another MCU in the transfer module is used to transmit the sampling data to the host computer through controller area network (CAN) bus. For the mechanical assembly (see Fig. 2), the WFT is directly mounted on the wheel center by two adapters via bolts, and it rotates with the spinning wheel. Usually, one adapter connects to the wheel hub and the other connects the modified rim. The transfer module is connected to the sample module by a pair of bearings, and they do not rotate. A lithium battery is used for the power supply, allowing the system to work continually over ten hours. All the modules are protected from water and dirt by the transducer casings. The above-mentioned system hardware and specifications can be also found in our recent work [4]",
            " Thus, the linear model of (1) for a two-axis WFT can be expressed as F = CV (2) where C is the 2 \u00d7 2 matrix that can be solved by least-square method or simply writing it with the pseudoinverse as C = FV + . (3) By encoding the model C into the host computer, the wheel force signal (unit: kN) can be computed from the output of the bridge (AD value). Another process is the rotational decoupling that converts the wheel force in wheel coordinate Ow Xw Y w Zw (local wframe) to the vehicle coordinate OvXvY vZv (vehicle v-frame) because the latter one is the actual definition of wheel loads in vehicle engineering (see Fig. 2). At the very beginning, the two coordinates coincide with each other. When the vehicle is in motion, the wheel coordinate will rotate around the y-axis, and their relationship is expressed as [29] \u23a1 \u23a3 xv yv zv \u23a4 \u23a6 = \u23a1 \u23a3 cos \u03b8 0 sin \u03b8 0 1 0 \u2212sin\u03b8 0 cos\u03b8 \u23a4 \u23a6 \u23a1 \u23a3 xw yw zw \u23a4 \u23a6 . (4) Note that the applied load f of (1) is in the local frame and it can be rewritten as fw = [fw x , fw z ]T . To get the actual forces (longitudinal force and vertical load) on the spinning wheel in the ground vehicle frame, f v is obtained from (4) as { fv x = fw x cos\u03b8 + fw z sin\u03b8 fv z = \u2212fw x sin\u03b8 + fw z cos\u03b8 (5) where the real-time rotational angle \u03b8 is detected by the encoder, and fw is computed by using the calibration matrix C (inbuilt parameters in host computer) in w-frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001216_demped.2019.8864903-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001216_demped.2019.8864903-Figure3-1.png",
        "caption": "Fig. 3. Motor\u2019s flux distribution under healthy and broken rotor bar conditions.",
        "texts": [
            " The FEM-based simulation of a three-phase induction motor, with the parameters shown in Table I, is performed under healthy, one, and two broken rotor bar conditions. Since the simulation is performed using 2D field analysis, the ignored end windings are compensated by adding additional resistances and inductances in series with coils. The per phase stator coils are series and parallel connections of copper strands, making the current density uniformly distributed. The simulation is performed at rated load under constant speed. The flux distribution under healthy and two broken rotor bar conditions is shown in Fig. 3. It is evident that the flux density increases across the broken bars, putting the adjacent bars under increased magnetic stress. The increase in the current of the neighboring bars makes the machine vulnerable to break more bars in time, if the fault is not timely diagnosed and repaired. The obtained results can be used as a benchmark to differentiate between harmonics due to motor itself (slot harmonics), due to a fault, such as broken rotor bars, and due to inverter. In Fig. 4, simulated three phase currents for the healthy case and frequency spectra for the healthy and the broken rotor bars is presented"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002937_012019-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002937_012019-Figure5-1.png",
        "caption": "Figure 5.  Fuzzy rule table output surface.",
        "texts": [
            " when the speed of mobile robot alternative space is larger, and the distance of target is relatively close, the less mobile machine environment barriers, simple and close to the target environment, mobile robot to the target point can choose to use a faster pace, so should choose the score function of weightings for  larger,  smaller,  larger. ACAE 2020 Journal of Physics: Conference Series 1905 (2021) 012019 IOP Publishing doi:10.1088/1742-6596/1905/1/012019 According to the above design ideas of fuzzy rules, the output surface of the input fuzzy rules is shown in Figure.5, the output surface of the input  fuzzy rules is shown in Figure 6, and the output surface of the input  fuzzy rules is shown in Figure.7. In order to verify the effectiveness of the algorithm in this paper, Windows10 operating system, Intel(R) core(TM) i5-8400 CPU@2.80GHz and memory 8G are adopted. Based on matlab 2016a, Then simulation planning of raster maps with different number of obstacles were established. In this experiment, the starting point position of mobile robot was set as (0,0), and the target point position was set as (7,6)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001438_lra.2020.2969161-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001438_lra.2020.2969161-Figure2-1.png",
        "caption": "Fig. 2. Concept and working principle of the proposed slave and master devices. (a) and (b-i) indicate the structural configuration of the forceps-driver and the working principle of the opening and closing motion, respectively. The force sensor module assembled on the rotator measures the force applied to the forceps-handle in contact with the sensor module. (c-i) and (d-i) show the structural configuration of the variable stiffness module consisting of six diaphragm flexures and two SMA wires. And, (e-1) and (e-2) schematically present the elastic deformation of the module that is caused by a user\u2019s pinching motion.",
        "texts": [
            " In this work, considering the essential functions that must be implemented in the forceps-driver and the master device of the tele-operated microsurgical robotic system, we aim to develop a novel sensor-embedded forceps-driver with high-precision gripping-force control and a master device with the function to display haptic feedback on the gripping-force. The conceptually designed both devices are fabricated as functional prototypes (see Fig. 1), and the working performances are experimentally investigated. Also, the practical feasibility and applicability are assessed through the user test performing the tele-operation utilizing the proposed system. The structural configuration and the working principle of the forceps-driver are schematically shown in Fig. 2(a) and (b-i), respectively. In designing the forceps-driver, we focus on implementing the device having the functions to drive commercial microsurgical forceps and to measure the gripping-force applied to the gripped-object by the forceps-tips. When a pair of microsurgical forceps is inserted into the driver and the forceps end is fixed by a clamp, the rotator is rotated by a servo motor connected via a tendon wire to press the forceps-handles slightly. In this motion, the force sensor module assembled on the rotator becomes into contact with one of forceps-handles, and the sensor begins to measure the force applied to the forceps-handle by the rotator. This initial state to drive the forceps is illustrated in Fig. 2(b-1). The end-tips of the forceps that are initially open are closed, when the rotator is rotated counterclockwise by the motor and clamps the forceps-handles accordingly (see Fig. 2(b-2)). In this operation, the angular displacement of the rotator is measured by a potentiometer configured on an upper cover. In developing the forceps-driver, we aim to implement the high-precision tiny modular force sensor suitable for the compact and lightweight forceps-driver. For this aim, we fabricate a force sensor module and embed into the device instead of using commercially available sensors. As explained the sensor fabrication process in [14], the sensor is implemented by casting the soft material (e",
            " The basic function required of the master device is to allow the surgeon to intuitively drive the forceps. To implement the function for the intuitive tele-operation, we consider the handoperation principle of the commercial forceps and attempt to design the driving mechanism according to the principle. Typically, the surgeon handles the forceps with the thumb and forefinger, and the forceps-tips are closed and/or opened by a surgeon\u2019s pinching motion. Based on this principle, we design a novel master device schematically shown in Fig. 2(c-1) and (c-2). The device has a structure in which the left part and the right part are mirror-symmetrical, and each part consists of diaphragm flexure elements arranged in multiple layers. In this structure, when an external force is applied to the push button that is the center of the diaphragm flexures by the pinching motion, the diaphragm flexures are elastically deformed as indicated in Fig. 2(e-1) and (e-2). The change in distance between the center of the left part and the center of the right part according to the pinching motion is measured by the magnetic displacement sensor that consists of a magnet embedded in the right part and a magnetic sensor PCB built in the left part. By designing the master device with this mechanism, the surgeon can tele-operate the forceps equipped on the slave-system with intuitive motion. Furthermore, because the intention to operate the forceps by the surgeon can be detected quantitatively by measuring the displacement varying according to pinching motion, the intention of the surgeon to operate the slave device can be directly reflected in controlling the device",
            " And, the Authorized licensed use limited to: University of Canberra. Downloaded on April 29,2020 at 07:28:20 UTC from IEEE Xplore. Restrictions apply. reaction force is perceived based on the proprioceptive sense in the muscles, joints, and tendons of the fingers. In this work, by taking into account this principle, we attempt to design the master device as a variable stiffness module that can apply variable reaction force to the fingers according to the gripping-force of the forceps. As illustrated in Fig. 2(d-1) and (d-2), a shape memory alloy (SMA) wire wound around pulleys is configured inside both the left part and the right part. In the case that the master device is implemented with this configuration, the effective stiffness of the module can be varied by the SMA wire that exhibits different stiffness depending on the activation state. The working principle of the proposed variable stiffness module can be explained clearly with Fig. 3. The half-model of the module is expressed as a simplified two-spring model as illustrated in Fig",
            " Even though Authorized licensed use limited to: University of Canberra. Downloaded on April 29,2020 at 07:28:20 UTC from IEEE Xplore. Restrictions apply. the slight error is observed between the data sets, the graph shows that the embedded force sensor module can function to measure the gripping-force appropriately. Also, the result demonstrates that the gripping-force measurement capacity of the device is about 4 N. Because the force sensor module applied to the proposed forceps-driver is configured to contact the forceps-handles as indicated in Fig. 2(a), the practical gripping-force applied to the gripped-object by the forceps-tips cannot be measured directly. For this reason, we consider the scheme to estimate the grippingforce indirectly by utilizing the embedded force sensor module and the potentiometer. The schematic diagram illustrated in Fig. 7(a) shows the gripping-force estimation scheme. Two lines in the graph indicate the measured force as a function of the angular displacement measured by the potentiometer during the motion that the open forceps-tips are closed by the rotation of the rotator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001594_j.cma.2020.112996-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001594_j.cma.2020.112996-Figure3-1.png",
        "caption": "Fig. 3. Quarter-circular arch clamped at one end and subjected to a transversal force at the free end.",
        "texts": [
            " Different polynomial orders are considered to examine the effect of continuity of the NURBS functions on the performance of the isogeometric formulations. The characters \u201cpn\u201d appearing after indicator \u201cR\u201d of the formulation title represent the polynomial orders used, e.g., \u201c\u03c6Rp3d0\u201d means a \u03c6Rd0 formulation with cubic NURBS functions. The classic Gauss\u2013Legendre quadrature rule with p + 1 integration points was performed for the numerical integration. Consider a quarter-circular arch clamped at one end and subjected to a transversal force P = [0 0 Pz]T at the free end, as in Fig. 3. The rod\u2019s axis lying on the x-y plane of the Cartesian system is r(\u03b2) = [\u2212\u03c1 cos \u03b2 \u03c1 sin \u03b2 0]T (71) with radius \u03c1 = 1.0 m, angle \u03b2 \u2208 [0, \u03c0/2], and the transversal vectors defined as: constant d2 = ez , d3 = d1 \u00d7 d2. We solved this problem under a unit transversal load Pz = 1kN, with the following material properties: E = 1.999 \u00d7 108 kN/m2, G = 7.996 \u00d7 107 kN/m2, and the following sectional properties: A = \u03c0r2, J = \u03c0r4/2, I2 = I3 = \u03c0r4/4 (a circular section with radius r = 0.01 m). For this problem, the exact vertical displacement uz (\u03b2) and angle of twist \u03c6(\u03b2) are uz (\u03b2) = Pz\u03c1 3 ( 2\u03b2 \u2212 sin \u03b2 \u2212 \u03b2 cos \u03b2 2G J + sin \u03b2 \u2212 \u03b2 cos \u03b2 2E I3 ) \u03c6(\u03b2) = Pz\u03c1 2 ( \u22123 sin \u03b2 \u2212 \u03b2 cos \u03b2 2G J + sin \u03b2 \u2212 \u03b2 cos \u03b2 2E I3 ) (72a,b) To examine the dependence of the invariance properties on the axis geometry for the isogeometric formulations concerned (especially the \u03c6Rd0 formulation), we considered additionally the B-spline interpolation for the rod\u2019s geometry using the collocation method [14], in comparison with the NURBS interpolation by which the circular axis is exactly represented"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001737_tim.2020.3003361-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001737_tim.2020.3003361-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of (a) regenerative effect and (b) cutting model.",
        "texts": [
            " In general, self-excited vibration includes speed feedback vibration, mode-coupling vibration and regenerative vibration. Among them, regenerative vibration is the main type of self-excited vibration in the cutting process. It will not only lead to the decrease of machining quality and increase of machining cost, but also will lead to the hidden danger in the process of machining. For online monitoring purpose, it is necessary to comprehend the signal feature of regenerative chatter. Regenerative effect is caused by the phase difference between the two sequential cutting tracks left by the tool as shown in Fig. 1(a). Fig. 1 (b) shows the cutting model of regeneration chatter during turning. Authorized licensed use limited to: University of Southern Queensland. Downloaded on July 05,2020 at 05:12:24 UTC from IEEE Xplore. Restrictions apply. 0018-9456 (c) 2020 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. 3 The initial relative vibration of the tool is caused by the unexpected disturbance during the cutting process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure46-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure46-1.png",
        "caption": "Fig. 46. Interrupted Core Ring",
        "texts": [
            " ( e ) The effort required to engage and disengage the synchro-coupling is a tiny fraction of the torque of the driving machine, and hence no servo-mechanism is required for its operation. The locking ring is light to operate and is entirely free from torque when it is moved into and out of engagement, hence a light actuating force only is required. 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 152 PROBLEMS OF FLUID COUPLINGS Reversing Gear. In the layout shown by Fig. 43 an independent reversing gear is introduced, of the kind shown diagrammatically by Fig. 46 ; hence the locomotive can be operated up to full speed in either direction. The synchro-couplings used in such a reversing gear would be of the controlled-pawl variety with a positive interlock to prevent the pawl control being operated except when the locomotive is stationary. If a bevel gear final drive to the axles could have been used the reverse gear might conveniently have been of the conventional type having two crown wheels and a double dog clutch, which can be moved to engage one or the other when driving in forward or reverse"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001209_jfm.2019.681-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001209_jfm.2019.681-Figure10-1.png",
        "caption": "FIGURE 10. (Colour online) Spatial distributions of the normalized exceptional maximum stored energy density Eexc of the second class of vibration in a 40 nm isotropic gold nanosphere and in the surrounding glycerol\u2013water mixture for the breathing mode n= 0 (a\u2013c) and quadrupolar mode n= 2 (d\u2013f ) using the compressional non-Newtonian model: (a,d) \u03c7 = 0 (pure water); (b,e) \u03c7 = 0.56; and (c, f ) \u03c7 = 0.85. The colour scale represents the exceptional maximum stored energy density normalized by the energy density at the north pole of the nanosphere for each mode and mass fraction. The radial thickness of the fluid computational domain is limited within 2b.",
        "texts": [
            " pure water), indicating that, for higher-order angular modes or smaller spherical radii with high vibration frequencies, the viscoelastic effect on fluid\u2013solid coupled vibrations becomes significant even in a low-viscosity fluid. As a consequence, the FSI vibration characteristics at nanoscale can be exploited to measure the viscoelastic properties of low-viscosity fluids (such as pure water) with faster relaxation times. In order to evidently demonstrate the stored energy evolution with the glycerol mass fraction, the spatial distributions of Eexc in a 40 nm isotropic nanosphere and different surrounding mixtures are illustrated in figure 10 for the breathing ht tp s: // do i.o rg /1 0. 10 17 /jf m .2 01 9. 68 1 D ow nl oa de d fr om h tt ps :// w w w .c am br id ge .o rg /c or e. A cc es s pa id b y th e U CS F Li br ar y, o n 05 O ct 2 01 9 at 1 3: 11 :4 6, s ub je ct to th e Ca m br id ge C or e te rm s of u se , a va ila bl e at h tt ps :// w w w .c am br id ge .o rg /c or e/ te rm s. (n = 0) and quadrupolar (n = 2) modes of the second class of vibration using the compressional non-Newtonian model. Note that for axisymmetric modes (i.e. m= 0), it is only necessary to give the energy distribution in the r\u2013\u03b8 plane. For the breathing mode, the energy distribution does not change along the \u03b8 -direction, while for the quadrupolar mode, the displacement and stress components depend on \u03b8 , resulting in the variation of the energy distribution with \u03b8 . However, the energy distribution of the quadrupolar mode is still symmetric, as shown in figure 10(d\u2013f ). It can be seen that, as expected, more energy is stored in the surrounding fluid due to the increasing elastic effect of the fluid as the glycerol mass fraction increases, which can compensate or even exceed the dissipated energy. The overall stiffness of the FSI system also increases with increasing glycerol mass fraction, which ultimately leads to the increase in the vibration frequency (not shown here for simplicity). For different glycerol mass fractions and different vibration modes, the variations of the quality factor of the second classes of vibration with the logarithmic spherical radius log10[b (nm)] are depicted in figure 11 using the Newtonian and compressional non-Newtonian fluid models"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure8-1.png",
        "caption": "Fig. 8. (a) The u-coordinate defined along the line of action on the transverse plane; (b) the \u2018contact plane\u2019 in the u-z coordinate, and three examples of possible configurations of lines of action, in different colors.",
        "texts": [
            " The contact force model presented in this section takes into account the general case of the load sharing and distribution for the helical gear meshing, such that the profile contact ratio can be an arbitrary value, not necessarily close to unity or falling into a certain interval, and there can be multiple simultaneous meshing lines of action (as in the case of the reference pump). Although this analysis can be generalized to non-involute conditions without significant modification, in the following the scope is limited to involute gearing, which works for the reference pump that uses an involute profile as the meshing profile. As shown in Fig. 8a, with the direction of the line of action on the transverse plane denoted as u, cutting along the blue line in Fig. 8a into the paper gives a cross-section shown in Fig. 8b, which is a plane with coordinate u and z. With the pitch line (the line parallel to both gear rotation axes, and with equal distance to each axis, u = 0 in Fig. 8) as the reference line, the interval u \u2208 [ \u2212 L/2, L/2] is the region with involute-to-involute contact. The length of L can be determined as: L = \u03b3 CRt = m\u03c0cos\u03b10CRt (29) in which CRt is the transverse contact ratio or profile contact ratio. The value of CRt is a function of various gear design parameters and the center distance between two gears. The detailed equations for involute gearing can be found in [5]. The change of profile contact ratio with varied center distance for the reference pump is shown in Fig",
            " Because of the linear relationship between the axial position and the phase shift of the transverse gear profile, the line of action in the u-z plane is a straight but inclined line. As the shift in u-direction between the contact point on the top cross-section and bottom cross-section is \u0398 \u22c5 rb, the angle \u03b8\u2019 between the line of action and the u-axis can be determined from the geometric relation: \u03b8\u2019 = arctan (\u0398 rb H ) = \u03b2b (30) which proves that the tilting angle of the helical line of action is equal to the helix angle on the base circle \u03b2b. In Fig. 8b, three examples of possible configurations of the line of actions are shown. Among them, the blue line of action is with its full contact length. While the other two, the yellow one and the green line, are incomplete contact, that is, the contact is not from the top to the bottom of the gear. In particular, the configuration like the green line only happens with the helical contact ratio is larger than the profile contact ratio. Assuming that the contact force Fc on the driving flank of the gear is only used to overcome the torque given by the fluid pressure on the driven gear (i.e. Gear 2), therefore the balance of the torque is written as Fc1cos\u03b2b mNcos\u03b10t 2 = TG2 (31) hence the magnitude of the driving contact force can be written as Fc1 = 2TG2 m N cos\u03b2bcos\u03b10t (32) transformed back to the Cartesian coordinate system, based on the coordinate settings illustrated in Fig. 8b: Fc1,x,G1 = \u2212 Fc1cos\u03b2bsin\u03b1t Fc1,y,G1 = Fc1cos\u03b2bcos\u03b1t Fc1,z,G1 = Fc1sin\u03b2b (33) while the contact force experienced by Gear 2 is of the same magnitude but the opposite direction, i.e. Fc1,x,G2 = Fc1cos\u03b2bsin\u03b1t Fc1,y,G2 = \u2212 Fc1cos\u03b2bcos\u03b1t Fc1,z,G2 = \u2212 Fc1sin\u03b2b (34) As shown in Fig. 10, for a particular contact line, its length Lc,i can be written as Lc,i = \u20d2 \u20d2 \u20d2 \u20d2 u1 \u2212 u2 sin\u03b2b \u20d2 \u20d2 \u20d2 \u20d2 (35) Two endpoints of the contact line are denoted as p1 and p2, respectively. From the relative position between p1, p2 and pmid as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002572_j.apm.2020.12.020-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002572_j.apm.2020.12.020-Figure9-1.png",
        "caption": "Fig. 9. Virtual prototype of 4-RSR&SS mechanism in ADAMS.",
        "texts": [
            " That is, the driving torques of the active joints obtained from the inverse rigid body dynamic model are used as the feed-forward input of the flexible multi-body dynamic model and the degenerated rigid body dynamic model of the mechanism, to solve the forward dynamic equations so as to obtain the dynamic response of the system. The dynamic response curves of generalized coordinates in the time domain are solved by substituting the torques into the dynamic model of flexible bodies. The driving torques are also applied to the virtual prototype of ADAMS shown in Fig. 9 to simulate the dynamic response of tracking mechanism. The theoretical dynamic response curves and ADAMS software simulation dynamic response curve are compared with each other to verify the effectiveness of the model. Using the Gill algorithm, the nonlinear time-varying ordinary differential elasto-dynamic equations are solved. In order to compare and analyze the dynamic performance of 4-RSR&SS parallel tracking mechanism in rigid multi-body dynamics and flexible multi-body dynamics, and to verify the correctness of the dynamic response analysis results of the model, 2 simulation examples are listed in this section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001250_j.triboint.2019.106028-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001250_j.triboint.2019.106028-Figure3-1.png",
        "caption": "Fig. 3. Contact model of two stepping tapered rollers in opposite orientation.",
        "texts": [
            " In improving the research efficiency, the instantaneous contact of the helical gears is approximately simulated using two tapered rollers with opposite orientations, and relevant studies have proven that the simplified model has a slight difference with the true contact state of the gear [35,36]. The working gear profile of the double involute gear is still involute, and the meshing principle is similar to those of common involute helical gears. However, after grading, the root is thickened, and the top is thinned. Hence, the instantaneous contact of double involute gears can be approximately simulated using two stepping tapered rollers in opposite orientation, as shown in Fig. 3. The teeth surface contact lines of the double involute gears are inclined and intermittent, and the contact line first changes from short to long, then from long to short, until it disengages from engagement. According to the relationship between the transverse ratio \u03b5\u03b1 and overlap ratio \u03b5\u03b2 of the double involute gears, the length of the contact Fig. 1. Basic tooth profile of double involute gears. Fig. 2. Diagram of transverse engagement of double involute gears. Z. Yin et al. Tribology International xxx (xxxx) xxx line from entering the contact area to completely exiting the meshing area can be divided into the following two cases, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure2-1.png",
        "caption": "Fig. 2 Schematic diagram of cutter coordinate system",
        "texts": [
            " When slicing the workpieces with internal teeth, the cutter is located inside the workpiece as shown in Fig. 1. There is an intersection angle \u03b3 between the cutter shaft and the workpiece shaft. The direction of the cutter tooth is required to be consistent with that of the workpiece slot. The cutter rotates with the workpiece at a certain speed ratio, and the workpiece feeds axially at the rate f. As the basis of mathematical derivation, the cutter coordinate system S2 (o2, x2, y2, z2) is first established as shown in Fig. 2, where axis z2 coincides with the centerline of the cutter. On the premise that the cutting edge is error-free, the rake face can be constructed in two ways: by a plane that is easy to manufacture or by a curved surface that can ensure consistent and reasonable working rake angle. The plane rake face is constructed in the following way to form rake angle. As shown in Fig. 3, vectors lbz, lbt, and lb are defined by the criterion of passing through the midpoint D (xD, yD, zD) of the addendum of the conjugate surface (see literature [6])"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003681_971510-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003681_971510-Figure5-1.png",
        "caption": "Figure 5. Finite element model of piston and fluid",
        "texts": [
            " The variation of contact surfaces are then determined by checking and computing the penetrations The finite element model of pads and backing plates consist of 756 solid elements and 1182 nodes as shown in Figure 3: The finite element of anchor braket consists of 639 solid elements and 1224 nodes as shown in Figure 4: The brake system consists of 5390 elements and 9368 nodes. FRICTION CURVES The coefficient of friction used in this case is a function of contact pressure and sliding velocity. Since the occurrence of brake squeal can be detected in a very short period of time, the variation of temperature has not been taken into consideration in this example. The friction data generated from the dynamometer test is shown in Figure 7. The pistons and fluid are modeled by using rods and springs respectively, as shown in Figure 5: RESULTS PISTON FORCES The time history of the force applied to each piston is shown in Figure 8. The time history of the acceleration at a selected node on the rotor is shown in Figure 9. ' f i e response is a combination of all the vibration modes of the rotor. In order to cornpare the results to test data, a tool using fast Fourier transformation (FFT) is (leveloped to compute the frequencies of the response. The distribution of the frequencies generated form EFT is shown in Figure 10. A plot showing the occurrences of the noise obtained from the experiments is also shown in Figure I I "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure5-1.png",
        "caption": "Fig. 5. Three Stages of Filling (Fan Drive)",
        "texts": [
            " Quantity-Slip Curves for Continuous Core Ring Coupling (Fan Drive) Fig. 3. Quantity-Slip Curves for Interrupted Core Ring Coupling (Fan Drive) cutting ports through, or interrupting sections of, the core guide ring at regular intervals, as shown in Fig. 4, Plate 1. The impeller and runner in the upper photograph have a standard \u201cVulcan\u2019) circuit with continuous core ring, and those below have the interrupted core ring. Three progressive degrees of filling of a Vulcan circuit which is oversized for the power to be transmitted are illustrated by Fig. 5, and the extent to which the core guide ring causes the \u201cflat spot\u2019) by interfering 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 87 with the progressive building up of the vortex ring is clear. The complete elimination of the core guide ring was naturally tried, but it is not a satisfactory solution because the slip, when full, is on the high side. Moreover the vortex tends to be unstable in form, and torque surges can occur under certain conditions of load and degree of filling"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002850_masy.202000307-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002850_masy.202000307-Figure1-1.png",
        "caption": "Figure 1. Shape and nominal sizes [mm] of the specimens.",
        "texts": [
            "8 mm\u2014named in the following SP08) were vertically manufactured (perpendicular to the start plate) by using the ARCAM A2X available at the Italian Aerospace Research Centre (CIRA). Standard process themes for net structure, developed by ARCAM AB for the Ti6Al4V Titanium alloy with a layer thickness equals to 50 \u00b5m, were used in this study. The process was carried out in vacuum (in the chamber, from 5 \u00d7 10\u22123 mbar at the start to 2 \u00d7 10\u22125 mbar at the end of the process). Shape and the other nominal sizes of the specimens are shown in Figure 1. Static tensile tests were carried out according to ASTME8 standard by electromechanical testing system INSTRON with a load cell capacity of 5 kN. Concerning the SP2 specimen, the strains were measured by means of an extensometer (Figure 2a) during the tests. Contrary, for SP08 specimen, due to the small diameter, it was not possible to use the extensometer. Hence, the strains have been calculated from the recorded crosshead displacements. The specimens were tested under displacement control with a crosshead speed v = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000745_0731684416678670-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000745_0731684416678670-Figure2-1.png",
        "caption": "Figure 2. Prepreg tow of non-geodesic fiber path achieved by in-plane bending deformation: (a) before fiber micro-buckled and (b) after fiber micro-buckled.",
        "texts": [
            " The length of fiber in one period of wave W can be obtained by following equation: W \u00bc 2 Z ds \u00bc 2 Z L 0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe \u00f0u0\u00de2 q dx \u00bc 2 Z L 0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe 2A L 2 cos2 2 x L s dx \u00f02\u00de Thus, the compressive strain of fiber in one period of wave can be expressed by equation (3). \" \u00bc l l \u00bc W L W \u00bc 1 L W \u00f03\u00de Combining with equation (2), equation (3) can be transformed into following equation by elliptic integral of the second kind and the Taylor expansion. \"\u00bc1 1 1\u00fe2 k2 8 \u00fe13 k2 8 2 \u00fe90 k2 8 3 \u00fe644 k2 8 4 \u00fe4708:5 k2 8 5 \u00fe . . . \u00f04\u00de where k2 \u00bc 2 A L 2 = 1\u00fe 2 A L 2 . at University of Otago Library on November 22, 2016jrp.sagepub.comDownloaded from On the other hand, the prepreg tow of non-geodesic fiber path achieved by in-plane bending deformation is shown in Figure 2. As shown in Figure 2, the fiber path of AFP process can be assumed to be the center line of prepreg tow in width direction. Thus, the length of center path l between central angle d can be calculated as equation (5), l \u00bc Rda \u00bc \u00f01=kg\u00deda \u00f05\u00de where R and kg are the geodesic curvature radius and geodesic curvature of fiber path, respectively. As the width of prepreg tow is w, the curvature radius of AD and BC are R+w/2 and R w/2, respectively. Therefore, the length of AD and BC can be expressed by equations (6) and (7)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure10-1.png",
        "caption": "Fig. 10. Two endpoints of a contact line p1 and p2 and their middle point pi as the equivalent acting point of the contact force of this contact line.",
        "texts": [
            " Gear 2), therefore the balance of the torque is written as Fc1cos\u03b2b mNcos\u03b10t 2 = TG2 (31) hence the magnitude of the driving contact force can be written as Fc1 = 2TG2 m N cos\u03b2bcos\u03b10t (32) transformed back to the Cartesian coordinate system, based on the coordinate settings illustrated in Fig. 8b: Fc1,x,G1 = \u2212 Fc1cos\u03b2bsin\u03b1t Fc1,y,G1 = Fc1cos\u03b2bcos\u03b1t Fc1,z,G1 = Fc1sin\u03b2b (33) while the contact force experienced by Gear 2 is of the same magnitude but the opposite direction, i.e. Fc1,x,G2 = Fc1cos\u03b2bsin\u03b1t Fc1,y,G2 = \u2212 Fc1cos\u03b2bcos\u03b1t Fc1,z,G2 = \u2212 Fc1sin\u03b2b (34) As shown in Fig. 10, for a particular contact line, its length Lc,i can be written as Lc,i = \u20d2 \u20d2 \u20d2 \u20d2 u1 \u2212 u2 sin\u03b2b \u20d2 \u20d2 \u20d2 \u20d2 (35) Two endpoints of the contact line are denoted as p1 and p2, respectively. From the relative position between p1, p2 and pmid as shown in Fig. 10, the coordinates of two endpoints can be determined as X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 u1i = min { max [( \u03d5 \u2212 \u03b4 \u2212 \u0398 2 ) rb, \u2212 L 2 ] , L 2 } u2i = min { max [( \u03d5 \u2212 \u03b4 + \u0398 2 ) rb, \u2212 L 2 ] , L 2 } (36) Another assumption to be made is that: the contact force is distributed uniformed by all the contact length, therefore for each contact line, with uniformly distributed load, the centroid of the load is at the center of the contact line, whose coordinates are ui = u1i + u2i 2 zi = z1i + z2i 2 (37) Based on this assumption, the total length of the contact line is the summation of all contact lines (Ncl contact lines assumed): Lc = \u2211Ncl i Lc,i (38) With the total contact force known from Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001779_lra.2020.3007467-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001779_lra.2020.3007467-Figure1-1.png",
        "caption": "Fig. 1. Partially constrained trajectories of the manipulation frame, e.g. \u2208 R3, leave uncertainties in grasp frame planning since the mobility of the mechanism is subject to constraints imposed by the closed kinematic chain. The proposed framework utilizes Model Predictive Control to solve for a valid grasp frame trajectory with any underconstrained reference.",
        "texts": [
            " Throughout this letter, we assume all hand and object motions are quasistatic and the weights of the objects used are negligible\u2013disregarding the need to explicitly model dynamics or object-specific properties, e.g. inertias. Moreover, we leverage a compliant end effector as these mechanisms are beneficial for maintaining stability of the hand-object system during manipulation, mitigating concerns of losing contact [7], [20]. The establishment of the grasp frame generalizes the geometric properties of an arbitrary object within a grasp [21]. Fundamentally, it portrays the local geometry of the object and standardizes the representation of the object frame (Fig. 1, 2). We will reference the object frame as being one in the same as the grasp frame, as we expect object weights to be negligible. Assuming a single non-rolling contact is maintained on each fingertip of a hand with k fingers, let us define contact points P = p1, . . . , pk where pi \u2208 R3,\u2200i \u2208 {1, . . . , k} with respect to the hand frame. Noteworthily, with non-rolling contacts, any 3 points in P can explicitly define the grasp frame. For simplicity, let\u2019s assume p1, p2, and p3 are used. Then, we can define the grasp frame pose, X \u2208 SE(3), by Gram-Schmidt orthogonalization, X = [Gx,Gy,Gz|Go] \u2208 SE(3) Go = 1 3 (p1 + p2 + p3) Gx = p2 \u2212 p1 ||p2 \u2212 p1||2 Gz = (p3 \u2212 p2)\u00d7 Gx ||(p3 \u2212 p2)\u00d7 Gx||2 Gy = Gz \u00d7 Gx (1) In this formulation, Gx,Gy , and Gz represent the directional vectors about the x, y, and z axes, respectively, with reference to the origin, Go"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003566_pime_proc_1945_153_010_02-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003566_pime_proc_1945_153_010_02-Figure15-1.png",
        "caption": "Fig. 15. Path of Coned Wheel Treads of Unequal Diameters",
        "texts": [
            "comDownloaded from 36 COMMUNICATIONS ON R I D I N G AND WEARING QUALITIES OF RAILWAY CARRIAGE TYRES the larger wheel would tend to drag. Had the authormy definite proof that small wheels always dragged? If the wheels were coned equally and the diameters were the same they would ride centrally on the circles, unless disturbed by a side blow, when lateral creep would take place on both wheels alternately. Should the wheels be of the same coning but of different diameters, they would tend to centralize with an inclined axle (see Fig. 15) with consequent flange wear. It seemed obvious that the wear of treads and flanges depended not only upon the original shape but also upon the restrictions imposed by the bogie. Here again there was a difference of opinion between various authorities; but was it not true that in the majortiy of cases wheels coned 1 in 20 wore to a 12-inch radius on the tread? He was struck by the five-figure dimensions of the wheels, which made him wonder whether the last figure was accurate, especially when taken on a coned wheel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002994_j.procir.2021.05.013-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002994_j.procir.2021.05.013-Figure10-1.png",
        "caption": "Fig. 10. Model of separated pliers consisting of five components.",
        "texts": [
            " Additionally, the volume of support structure can be reduced if a complex component is separated into several components, which are oriented independently for minimum support structure. On the other hand, components can be separated to use further potentials of AM, such as interchangeability of wear components, modularization, etc. In the case of the presented pliers, the geometry fits into the build space and the amount of support structure is low. Therefore, it is not necessary to separate the component due to manufacturing restrictions. Fig. 10. Model of separated pliers consisting of five components. According to the last guideline in Fig. 6 there are many kinds of pliers, as for example side cutters, needle-nose pliers, pincers, and crimping pliers. The difference between these is mainly the geometry of the jaws. By separating the jaws, it is possible to interchange the front component of the pliers and fulfill several functions with different attachments and the same base. Accordingly, one pair of pliers could be produced with different attachments, which is less expensive than a whole set of pliers",
            " Accordingly, one pair of pliers could be produced with different attachments, which is less expensive than a whole set of pliers. In addition, conventional production of the 84 Jannik Reichwein et al. / Procedia CIRP 100 (2021) 79\u201384 6 Author name / Procedia CIRP 00 (2019) 000 000 basic module of the pliers would be possible and additive production only of the attachments, which are required in much smaller quantities and can sometimes be very complex in terms of geometry (e.g., crimping pliers). The final product architecture is shown in Fig. 10 and the assembled design in Fig. 11. It was shown that by modelling the product, a restructuring of the product architecture is possible without being influenced by the original design, which is characterized by manufacturing restrictions of conventional processes. The final modular design allows a better use of the potentials of additive manufacturing, such as the possibility to manufacture complex geometries and already assembled joints. Additive manufacturing opens up new potentials in production due to great freedom in geometric design and costeffective production of individual components in small quantities"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000222_0954406219885979-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000222_0954406219885979-Figure4-1.png",
        "caption": "Figure 4. Force analysis for the ball in quasi-static equilibrium state.",
        "texts": [
            " According to the Pythagorean theorem, the displacement equations for the ball are as follows R2 bqj \u00fe Z2 bqj ro ur 0:5D\u00fe oqj 2 \u00bc 0 \u00f04\u00de Riqj Rbqj 2 \u00fe Ziqj Zbqj 2 ri 0:5D\u00fe iqj 2 \u00bc 0 \u00f05\u00de where iqj and oqj denote the inner and outer contact deformations respectively;Rbqj and Zbqj denote the radial and axial distance between the ball center and the outer raceway groove curvature center respectively; Riqj and Ziqj denote the radial and axial distance between the inner and outer raceway groove curvature centers respectively, which can be calculated by using the following equations Riqj \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2iqj \u00fe y2iqj q \u00f06\u00de Ziqj \u00bc ziqj \u00fe Li 2 \u00f07\u00de Load equilibrium for the ball and the inner ring The force analysis of the ball in quasi-static equilibrium state is shown in Figure 4. The constraint equations at xq yq zq 1 2 6664 3 7775 \u00bc cos y 0 1\u00f0 \u00deq sin y 1\u00f0 \u00deq Li 2 sin y \u00fe x 1\u00f0 \u00deq sin x sin y cos x sin x cos y 1\u00f0 \u00deq 1\u00f0 \u00deq\u00fe1 Li 2 sin x cos y \u00fe y 1\u00f0 \u00deq\u00fe1 cos x sin y sin x cos x cos y 1\u00f0 \u00deq 1\u00f0 \u00deq Li 2 cos x cos y 1 \u00fe z 0 0 0 1 2 66664 3 77775 x0q y0q z0q 1 2 6664 3 7775 \u00f02\u00de the ball are2,26 Qiqj sin iqj Qoqj sin oqj \u00fe 2Mgqj D cos oqj \u00bc 0 \u00f08\u00de Qiqj cos iqj Qoqj cos oqj 2Mgqj D sin oqj \u00fe Fcqj \u00bc 0 \u00f09\u00de where Qiqj and Qoqj denote the inner and outer contact forces respectively; Fcqj and Mgqj denote the centrifugal force and the gyroscopic moment at the ball; iqj and oqj denote the inner and outer contact angles respectively, which can be obtained from Figure 3 iqj \u00bc cos 1 Riqj Rbqj ri 0:5D\u00fe iqj \u00f010\u00de oqj \u00bc cos 1 Rbqj ro ur 0:5D\u00fe oqj \u00f011\u00de The contact forces can be calculated by using the Hertzian contact theory,2 and the centrifugal force and the gyroscopic moment can be expressed as follows Fcqj \u00bc 1 2 mdm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002880_s40998-021-00421-0-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002880_s40998-021-00421-0-Figure9-1.png",
        "caption": "Fig. 9 Structure chart of bulb tubular turbine",
        "texts": [
            " The classification accuracies of VMD-LDA-WNN for the four state signals are 100%, 97%, 98%, 100% respectively and the average accuracy is 98.75%. In contrast, the method proposed in this paper has more advantages in fault diagnosis. In this section, the vibration data of the near Wake Island Hydropower Station in Hunan province are used to validate the practicability and effectiveness of VMD-SVD in fault diagnosis. The unit of the near Wake Island Hydropower Station is bulb tubular turbine, the structure chart of bulb tubular turbine is shown in Fig. 9. The unit is composed of the light bulb head, main shaft, bearing, generator and so on. The main shaft is supported by two bearings, which bear mainly the radial and also certain axial load. An acceleration sensor CT1010 LC is mounted on bearing pedestal to acquire vibration signals and the sampling frequency is 10 kHz. Parameters of CT1010 LC are as follow: measuring range of 50 g (g \u00bc 9:8m/s2), sensitivity of 100 mV/g and frequency range of 0.5\u20135000 Hz. Three kinds of faults are simulated, such as sweep chamber, metal block friction and double bar percussion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000581_j.mprp.2016.04.087-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000581_j.mprp.2016.04.087-Figure4-1.png",
        "caption": "FIGURE 4 (Left) vertically and horizontally oriented tensile bars during SLM build-up; (right) horizontally manufactured tensile bar with different orientation of scanning vectors relative to the bar axis.",
        "texts": [
            " FIGURE 5 Tensile properties of SLM-processed 14Ni (200 grade) maraging steel in asbuilt and heat-treated (HT) conditions. S P E C IA L F E A T U R E significant differences, the precipitation heat treatment that substantially increases the hardness, tends to eliminate these differences. This effect was explained by the precipitation developing isotropically and acting as a dislocation barrier in all directions. SLM-processed test samples were manufactured in the same two orientations for measuring static mechanical properties. Figure 4 illustrates the differences between the bars: one set produced in an upright, vertical orientation and the other two samples oriented horizontally during the SLM process. For the latter two samples the orientation of the scanning vectors in the form of islands relative to the bar axis was divided into two groups. In the first horizontally scanned sample the vectors were parallel and perpendicular to the bar axis \u2013 referred to as 08/908 \u2013 whereas the directions changed to 458 for the second sample \u2013 referred to as 458"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003870_(sici)1097-0363(19990815)30:7<845::aid-fld867>3.0.co;2-o-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003870_(sici)1097-0363(19990815)30:7<845::aid-fld867>3.0.co;2-o-Figure2-1.png",
        "caption": "Figure 2. Hydrodynamic lubrication fluid element.",
        "texts": [
            " In general, full simultaneously converged field solutions for fluid film profile thickness, temperature, viscosity and pressure were regularly obtained for all but the severest of load conditions. By this means a great many predicted performance characteristics were studied. This paper outlines the design and development of the mathematical model and computation process itself and gives observations arising from its extensive use. The governing Reynolds\u2019 equation for the fluid film pressure field acting on one adjustable segment was derived with reference to the fluid element depicted in Figure 2. Pressure, film thickness and viscosity were set to vary in two dimensions; circumferentially around (x) and axially along (z) the segment surface. The derivation was based on a number of assumptions, most of which are normally invoked as follows: Fluid density constant (i.e. not affected by pressure or temperature). Fluid acceleration forces negligible compared with viscous shear forces. Weight of fluid is negligible. Film thickness small compared with radii of curvature. Pressure constant radially through film thickness at each point"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002559_j.mechmachtheory.2020.104218-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002559_j.mechmachtheory.2020.104218-Figure1-1.png",
        "caption": "Fig. 1. Illustration of a spherical four-bar linkage.",
        "texts": [
            " The rest of the paper is organized as follows. Section 2 defines the problem to be studied. A method of expressing position vectors with parameterized coordinates is introduced in Section 3 . With this method, algebraic coupler curve was derived in Section 4 . In Section 5 , we extend the study of coupler curve to curvature analysis and path synthesis. Examples are included in Section 6 . The work is discussed and concluded finally in Section 7 . 2. Problem formulation A spherical four-bar linkage is depicted in Fig. 1 , with its four linkage dimensions noted by { \u03b1 j }4 1 . The two grounded revolute joints are labelled B and D, the points at which their axes intersect the unit sphere. The two moving revolute joints are labelled A and C, the points at which their axes intersect the same sphere. The axes of the rotation of one dyad are thus given by the segments OB and OA ; the position vectors of B and A are b and a , both of unit magnitude, i.e., \u2016 b \u2016 = 1 , \u2016 a \u2016 = 1 (2) Likewise, the position vectors of points C and D are denoted by the unit vectors c and d ",
            " (3) can be expressed in a compact form as n 3 = Nn \u2032 3 (7) where N = [ n 1 , n 2 , n d ] ; n \u2032 3 = [ \u03bc, \u03bd, \u03bb] T (8) As we can see, matrix N contains a basis of R 3 , while vector n \u2032 3 includes three parameterized coordinates. Moreover, the vector of parameterized coordinates is an invariant, as it is frame independent. 4. Coupler curve of spherical four-bar linkages Now we can derive explicitly the coupler curve equation for spherical four-bar linkages. Referring to the spherical four- bar linkage shown in Fig. 1 , vector p = [ x, y, z] T of the coupler point P is parallel to OP and has a unit length, or g(x, y, z) = x 2 + y 2 + z 2 \u2212 1 = 0 (9) . We start from the coupler link AP C, for which the angle between OA and OP remains constant. An equation is thus obtained upon imposing this geometric constraint, i.e., p T a = cos \u03c11 (10) where a = Ra 0 (11) with a 0 being the unit vector parallel to the initial position of OA . Moreover, matrix R describes the rotation of link BA about OB . Using natural invariants of the rigid-body rotation, matrix R for a rotation about axis parallel to a unit vector e about an angle \u03d5 takes the form [24] R = ee T + s \u0303 e + c(1 \u2212 ee T ) , c \u2261 cos \u03d5, s \u2261 sin \u03d5 (12) where the sign \u2019 \u223c\u2019 on top of a vector denotes the anti-skew symmetric matrix derived from the vector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003709_s0043-1648(96)07486-8-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003709_s0043-1648(96)07486-8-Figure10-1.png",
        "caption": "Fig. 10. Photograph and schematic of the test rig, courtesy of Ha\u0308gglunds Drives AB.",
        "texts": [
            " Both the stick and slip case and the Coulomb friction case are studied. Figs. 8 and 9 show that there are large differences in the tangential load distribution and small differences in the location of the zero-sliding points for the two test cases. To be able to study how the wear depends on the number of revolutions (sliding distance), ten tests were performed Journal: WEA (Wear) Article: 7486 (see Table 2). The test rig can be seen as two hydraulic motors (Ha\u0308gglunds Drives AB Marathon motor) in parallel (see Fig. 10). Due to the construction of the test rig, two bearings were tested each time. Each test was performedwith standard bearings (SKF 29416 E), under constant load and rotational speed and using the same oil. The basic dynamic load rating for the bearing is about five times the applied axial load. To ensure a constant viscosity, the oil temperature was held constant at 50 8C. The hydraulic oil was filtered through a 10 mm filter. Running parameters were: load, 113 kN; rotational speed, 10 rpm; lubricant, mineral oil Shell Tellus 68 S; specific film thickness, lf0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001535_j.ymssp.2020.106723-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001535_j.ymssp.2020.106723-Figure3-1.png",
        "caption": "Fig. 3. Schematic of cardan shaft with misalignment in X-Y-Z axes.",
        "texts": [
            " The sleeve bearing rotates around the cross-shaft neck through needles. The structure of a universal joint used in cardan shaft and detail structure in sleeve bearing are shown in Fig. 1, respectively Two typical failures of the cardan shaft are abnormal wear of wear pad in sleeve bearing as shown in Fig. 2(a) and radial runout of cardan shaft as shown in Fig. 2(b). Due to the abnormal wear of sleeve bearing, it will result in excessive clearance which leads to shaft misalignment. An illustration of cardan shaft with misalignment in X-Y-Z axes is shown in Fig. 3. According to the installation of cardan shaft in CRH train, the drive shaft and driven shaft are set parallel with each other but in different planes. As shown in Fig. 3, cardan shaft misalignment is due to the offsets on cross-shaft. The offsets are named h1 and h2 at the drive shaft end, l1 and l2 at the driven shaft end. The offset h1 is defined as the shift distance at cross-shaft connected with the drive shaft relative to the drive shaft joint fork. The offset h2 and l2 is defined as the shift distance at cross-shaft connected with the connect shaft relative to center of the cross-shaft. The offset l1 is defined as the shift distance at cross-shaft connected with the driven shaft relative to the driven shaft joint fork",
            " The symbol O in the figure refers to the intersection of drive shaft axle and cross-shaft; symbol A refers to the center of the cross-shaft at drive side; and symbol B refers to the intersection point of cross-shaft at the drive side and connect shaft. The symbol D refers to the center of cross-shaft at driven side; symbol C refers to the intersection point of cross-shaft at the driven side and connect shaft; and symbol E refers to the intersection of driven shaft axle and crossshaft. To derive the kinematic equations of cardan shaft, an absolute coordinate system is built. The axes are marked X ,Y, and Z in Fig. 3. The origin of the coordinate is set at O, the X-axis is along with the drive shaft axis, the Y-axis is parallel with the direction at initial direction of line AO and Z-axis is perpendicular with the directions of X-axis and Y-axis. The coordinate of point E indicating the position of driven shaft is relative to drive shaft and is fixed in the model, which is XE YE ZE\u00f0 \u00de. Line AB is perpendicular to line AO and line BC. The existence of offsets h2 and l2 forced line AB to rotate around line AO with a certain angle g"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001966_j.mechmachtheory.2020.104136-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001966_j.mechmachtheory.2020.104136-Figure11-1.png",
        "caption": "Fig. 11. Equivalent model of driveline system for vehicle start-up.",
        "texts": [
            " k ( j ) n 1 k ( j ) nn \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 C m + n \u22121 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 c ( i ) 11 \u00b7 \u00b7 \u00b7 c ( i ) mm . . . . . . . . . c ( i ) m 1 \u00b7 \u00b7 \u00b7 c ( i ) mm + c ( j ) 11 c ( j ) 12 \u00b7 \u00b7 \u00b7 c ( j ) 1 n c ( j ) 21 . . . . . . . . . c ( j ) n 1 c ( j ) nn \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 Where, the subscripts in the parentheses denote the basic modules. With the basic model of engine, clutch, gear box and the resistance, the dynamic model of the driveline can be combined, using the method described above. The equivalent model of the driveline for start-up can be illustrated as shown in Fig. 11 . The dynamic model of this driveline can be expressed as Eq. (14) . J \u0308X + C \u02d9 X + KX = T (14) where, J = diag([ J1 , J2 , J3 , J4 , J5 , J6 , J7 , J8 ]), X = [ \u03b8 1, \u03b8 2, \u03b8 3, \u03b8 4, \u03b8 5, \u03b8 6, \u03b8 7, \u03b8 8] K = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 k 12 \u2212k 12 \u2212k 12 k 12 k 34 \u2212k 34 \u2212k 34 k 34 + k 45 r 2 4 \u2212k 45 r 4 r 5 \u2212k 45 r 4 r 5 k 45 r 2 5 + k 56 \u2212k 56 \u2212k 56 k 67 r 2 6 + k 56 \u2212k 67 r 6 r 7 \u2212k 67 r 6 r 7 k 67 r 2 7 + k 78 \u2212k 78 \u2212k 78 k 78 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 C = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 c 12 \u2212c 12 \u2212c 12 c 12 + kzR R \u2032 F \u2212kzR R \u2032 F \u2212kzR R \u2032 F c 34 + kzR R \u2032 F \u2212c 34 \u2212c 34 c 34 + c 45 r 2 4 \u2212c 45 r 4 r 5 \u2212c 45 r 4 r 5 c 56 + c 45 r 2 5 \u2212c 56 \u2212c 56 c 56 + c 67 r 2 6 \u2212c 67 r 6 r 7 \u2212c 67 r 6 r 7 c 8 + c 67 r 2 7 \u2212c 78 \u2212c 78 c 78 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 T = [ T e \u2212\u03bcs zRF \u03bcs zRF 0 0 0 0 \u2212T L ] where, T e is the output torque of the engine, T L is the resistive torque, J 1 is the inertia of the flywheel, J 2 is the inertia of clutch drum, J 3 is the inertia of clutch hub, J 4 is the inertia of driving gear of first gear set and all that is attached to it, J 5 is the inertia of driven gear of the first gear set and all that is attached to it, J 6 is the inertia of driving gear of the reducer gear set and all that is attached to it, J 7 is the inertia of driven gear of the reducer gear set and all that is attached to it, and J 8 is the equivalent inertia of the whole vehicle; k 12 is the engine output shaft stiffness, k 34 is the stiffness of input shaft of gear box, k 45 is the mesh stiffness of the first gear set, k 56 is the stiffness of the intermediate shaft in the gear box, k 67 is the mesh stiffness of the reducer gear set, and k 78 is the stiffness of the driving axle; c 12 is the equivalent damping of engine output shaft, c 34 is the equivalent damping of input shaft of gear box, c 45 is the mesh damping of the first gear set, c 56 is the equivalent damping of the intermediate shaft in the gear box, c 67 is the mesh damping of the reducer gear set, and c 78 is the equivalent damping of the driving axle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001319_s40430-019-2073-4-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001319_s40430-019-2073-4-Figure7-1.png",
        "caption": "Fig. 7 Path of circular motion (dimensions are in meters)",
        "texts": [],
        "surrounding_texts": [
            "In this part, a simulation in MATLAB environment was performed to evaluate the efficiency of the fuzzy sliding mode control. To this end, a planar robot with two degrees of freedom (Fig.\u00a01) was constructed. As shown in Fig.\u00a05, such as a real robot, noise, disturbance, mass uncertainty, and sampling discontinuity were also taken into account in the model. In this model, masses of the robot change as described in Sect.\u00a03 and some disturbances are added to the robot as shown in Eqs.\u00a031 and 32. Then, control methods were applied to the model and compared with each other. The block diagram of the control system that is simulated in Simulink is shown in Fig.\u00a06. According to Figs.\u00a07 and 8, the paths of interest for the model are a circular path and a linear path. For each one of these paths, the optimal classical sliding mode control and the optimal fuzzy sliding mode control are designed, and then their results are compared. Also, the end effector of the robot goes from point A to B and vice versa in 5\u00a0s. q1d, q2d, q\u03071d, q\u03072d, q\u03081d, q\u03082d can be calculated using reverse kinematics relations. As mentioned, the coefficients k1, k2, 1, 2, 1, 2 and k2 (sliding mode control), and the coeff icients k1, k2, 1, 2, a11, a12, a13, a21, a22, a23, b11, b12, b13, b21 , b22, b23 (fuzzy sliding mode control) (27)Scircular = R 2 ( 1 \u2212 cos ( 2 t 5 )) (28)Slinear = 0.8 2 ( 1 \u2212 cos ( 2 t 5 )) = 0.4 ( 1 \u2212 cos ( 2 t 5 )) are determined by the operator. Some properties of the improved system may be that the function gets worse as the coefficients tend to change. Therefore, since there is a contradiction concerning the optimization of the coefficients Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:5 1 3 Page 7 of 12 5 between the desired functions, a multi-objective optimization should be performed in order to select the coefficients. In this article, multi-objective genetic algorithm (NSGA II) was used to optimize the problem. Objective functions are required to apply genetic algorithm. In this article, it has been used with the aim of reducing errors and inputs of double control functions, which are as follows: 1. The function, which has the norm sum of the path error, was used to reduce tracking error, as shown below: The variables xi and yi are the coordinates of end effector in time ti . The variables xdi and ydi are the coordinates that the robot must cross during time ti 2. The following function was used to reduce the control inputs: The parameters 1i and 2i are the inputs of controllers for junctions 1 and 2 in time ti . The timing step is 0.01\u00a0s, and the functions fitness1 and fitness2 are computed for the duration of 5\u00a0s. (29) fitness1 = \u2016\u2016 \ufffd\ufffd\u20d7e1\u2016\u2016 + \u2016\u2016 \ufffd\ufffd\u20d7e2\u2016\u2016 = \u221a\u221a\u221a\u221a n\u2211 i=1 ( xi \u2212 xdi )2 + \u221a\u221a\u221a\u221a n\u2211 i=1 ( yi \u2212 ydi )2 (30) finess2 = \u2016\u2016 \ufffd\ufffd\u20d7\ud835\udf0f1\u2016\u2016 + \u2016\u2016 \ufffd\ufffd\u20d7\ud835\udf0f2\u2016\u2016 = \u221a\u221a\u221a\u221a n\u2211 i=1 ( \ud835\udf0f1i )2 + \u221a\u221a\u221a\u221a n\u2211 i=1 ( \ud835\udf0f2i )2 The following table shows the important components of genetic algorithm to implement the program. The ranges of designed variables of the sliding mode control for each path of interest have been considered as follows: The ranges of designed variables for the two paths in fuzzy sliding mode control were considered as follows: Genetic algorithm was performed several times, and during each performance, the primary population differed from the other performances. The total of the final answers are derived from the answers of these performances. This set of answers includes optimal points, none of which take priority over each other. In other words, one cannot find any member of the set that has priority over another member of the same set, as far as objective functions of both are concerned. The following chart shows the Pareto resulting from optimization. Each point of the chart represents the term for path tracking error and the term for controller inputs. According to Pareto plots (Figs.\u00a09, 10, 11, 12), it can be concluded that there are less controller inputs in control of fuzzy sliding mode as compared to the control of classical t\u20d7 = [0, 0.01, 0.02, 0.03,\u2026 , 5] 5 \u2264 k1, k2, k3 \u2264 200 1 \u2264 1, 2, 3 \u2264 200 0.1 \u2264 1, 2, 3 \u2264 50 5 \u2264 k1, k2, k3 \u2264 200 1 \u2264 1, 2, 3 \u2264 200 0.1 \u2264 a11, a12, a13, a21, a22, a23 \u2264 50 0.1 \u2264 b11, b12, b13, b21, b22, b23 \u2264 50 Fig. 8 Path of linear motion (dimensions are in meters) Fig. 9 Pareto chart resulting from optimization\u2013SMC\u2013circular path Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:5 1 3 5 Page 8 of 12 sliding mode. On the other hand, the amount of tracking error shows considerable decrease in the case of control of fuzzy sliding mode. In both controllers (fuzzy and classic), each point of Pareto plots includes unique parameters, as shown below: The points A, B, C are the best appropriate points to design the objective functions, although all obtained points are optimal and each one can be selected as an option for the designers. This means if the mono-objective error of the path k1, k2, 1, 2, 1, 2, k1, k2, 1, 2, a11, a12, a13, a21, a22, a23, b11, b12, b13, , b21, b22, b23 was being optimized, then point A was reached, whereas if the mono-objective of the controller input was being optimized, then point C was obtained. Comparing the points on the chart shown in Fig.\u00a09, B, also known as the break point, was selected as the best answer. The main property of this point is that a significant reduction can be seen not only in input torque but also in the tracking error of the path. The following chart shows the path error and input controllers for A, B, and C to provide a better understanding. The amount of designing variables and the values taken by objective function for point B are represented in Tables\u00a05, 6, 7, and 8. Since the disturbance vector was considered during optimization, classical sliding mode control and fuzzy sliding mode control responded appropriately to the perturbation vector. Hence, the amount and time of turbulence entered into the system were increased: W\u20d7(q\u20d7, \u20d7\u0307q, \u20d7\u0308q) = { W1 W2 } (31)W1 = { \u221280(N \u22c5 m) 1.5 \u2264 t \u2264 2 0 otherwise Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:5 1 3 Page 9 of 12 5 By selecting point B for the variables and applying it in the program, the following results were obtained. For better comparison, the methods of both classical sliding mode and fuzzy sliding mode are shown in these figures."
        ]
    },
    {
        "image_filename": "designv11_14_0003541_s0021859600045585-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003541_s0021859600045585-Figure5-1.png",
        "caption": "Fig. 5.",
        "texts": [
            " 4 give the work of separation obtained by graphical integration of these results and the theoretical curve. It will be seen that though the agreement for the total work is not very close the general slope of the relation is concordant: the force is much as calculated up to the point of drop rupture, but this occurs earlier than the simple geometrical assumptions suggest. This sidered as fixed. There is a downward force P due to the capillary attraction, which decreases as the separation 2x between the balls increases, and can be represented as in Fig. 5B by some curve whose general slope will be as the line P. There is an upward force Q exerted by the spring beam (in excess of that to balance the weight R of the upper ball), which also decreases as 2x increases, and can 138 On the capillary forces in an idealized soil (since Hooke's law will be obeyed at least for small displacements) be represented by a straight line whose general direction of slope is as for line Qt in Fig. 5B. With such a system, the upper ball force-displacement is represented by a line as Q% in Fig. 5 B the system will be unstable, for an increase in 2x will mean that Q is now greater than P, and the upper ball will continue to move up with, in will take up a position of stable equilibrium (P = Q) at some separation 2xe: since if 2x increases above this value, P becomes greater than Q. However, if the sensitivity of the balance arm is such that its general, resulting rupture of the drop. Accordingly, it is necessary that the sensitivity of the balance (displacement . , \u2014 1 should be less than that given by force / E"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001535_j.ymssp.2020.106723-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001535_j.ymssp.2020.106723-Figure1-1.png",
        "caption": "Fig. 1. Structure of the universal joint.",
        "texts": [
            " A typical constant velocity cardan shaft system consists drive shaft and driven shaft which are connected by a cardan shaft. The length of cardan shaft is variable depending on the motion of the shaft and can adapt different relative angles between the drive and driven shafts. The universal joint of cardan shaft contains cross-shaft and sleeve bearing. The sleeve bearing rotates around the cross-shaft neck through needles. The structure of a universal joint used in cardan shaft and detail structure in sleeve bearing are shown in Fig. 1, respectively Two typical failures of the cardan shaft are abnormal wear of wear pad in sleeve bearing as shown in Fig. 2(a) and radial runout of cardan shaft as shown in Fig. 2(b). Due to the abnormal wear of sleeve bearing, it will result in excessive clearance which leads to shaft misalignment. An illustration of cardan shaft with misalignment in X-Y-Z axes is shown in Fig. 3. According to the installation of cardan shaft in CRH train, the drive shaft and driven shaft are set parallel with each other but in different planes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure4-1.png",
        "caption": "Fig. 4 Cutting angle system and rake angle",
        "texts": [
            " Equation (8) is solved by Newton iterative method, and the unknown parameters are obtained, expressed as \u03b8i, \u03bai, and \u03bbi, correspondingly to every discrete value \u03bci. By taking \u03bai and \u03bbi into Eq. (7), a point on the cutting edge can be obtained as rdr \u00bc xp j \u03bai; \u03bbi\u00f0 \u00dei2\u00feyp j \u03bai; \u03bbi\u00f0 \u00de j2 \u00fe zp j \u03bai; \u03bbi\u00f0 \u00dek2 \u00f09\u00de A certain number of discrete points can be obtained in the same way, and the cutting edge is constructed using cubic Bspline curve by these points. Then, the tangent vector tr of the cutting edge can be obtained. The rake face is constructed taking the cutting edge as the boundary. As shown in Fig. 4, a coordinate system for cutting angle is defined at point M of the cutting edge according to the cutting speed v21 and the tangent vector tr. Among them, plane Ps is the cutting plane which passes through v21 and tr. Plane Pr is the base plane which is perpendicular to v21. Plane Po is the orthogonal plane which is perpendicular to Ps and Pr. According to metal cutting theory, the working rake angle is defined as the intersection angle between two straight lines in plane Po. One is the tangent line of the intersection curve of plane Po and rake face, and the other one is the intersection line of plane Po and plane Pr"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001158_rpj-09-2018-0227-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001158_rpj-09-2018-0227-Figure3-1.png",
        "caption": "Figure 3 The design of the roller module",
        "texts": [
            " Roller-type of leveling mechanism was chosen for MJ process, since it can remove excess material and flatten the surface at the same time. The dimension of roller module was limited to the space available within the printer and was also referred to the design in the commercial MJ systems. To simplify the design, there was only one UV light module to cure jetted droplets. The roller module was placed between printhead module and light source module along the y-axis, as shown in Figure 1. The design of roller module is shown in Figure 3. The roller is driven by a stepper motor next to it connected by a belt. There is a waste collecting tank with a scraper on the side that would scrape the excess resin from the roller surface. The upper plate for attaching the module to the x-axis can adjust the height of the roller module on both ends. The length of leveling region in x-axis was set to be 67mm to fully cover the nozzle region, 65mm, of printhead. The overall length of roller was 105mm. Moreover, the contact area between roller and material to be removed is proportional to the roller diameter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure5-1.png",
        "caption": "Fig. 5. Configurations of Mechanism I in circuit 2: (a) Configuration D: h1 = \u2212p/2, (b) configuration E: h1 = 0, and (c) configuration F: h1 = p/2.",
        "texts": [
            " It is noted that joints 1 and 6 in this prototype are prevented from full-cycle rotation due to interference between links 2 and 4 as well as links 1 and 5. Let K1, K2 and K3 denote the intersections of joint axes of joints 1 and 6, joints 2 and 5, and joints 3 and 4. P1, P2 and P3 represent the plane defined by the axes of joints 1 and 6, joints 2 and 5, and joints 3 and 4 respectively (Fig. 7). From Ref. [36], we obtain that planes P1,P2 and P3 and plane K1K2K3 have a common point, K, at any configuration of the 6R mechanism during motion. It is noted that at configurations A, C (Fig. 4), D and F (Fig. 5) of Mechanism I, point K2 is at infinity. A 6R mechanism that has three pairs of R joints with intersecting joint axes has been proposed using a geometric construction approach. Kinematic analysis of the mechanism has been presented. The analysis has shown that the 6R usually has two solutions to the kinematic analysis for a given input. In two configurations in each circuit of the 6R mechanism, the axes of four R joints are coplanar, and the axes of the other two R joints are perpendicular to the plane defined by the above four R joints"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001752_j.mechmachtheory.2020.103992-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001752_j.mechmachtheory.2020.103992-Figure6-1.png",
        "caption": "Fig. 6. Contact ellipse between (1) and (2) .",
        "texts": [
            " (57) can be written as K (12) = K (12 u ) + K (12 t) 2 + sin ( 2 + 2 0 ) \u221a ( K (12 u ) \u2212 K (12 t) 2 )2 + (\u03c4 (12 t) ) 2 , (58) where 0 = 1 2 arctan ( K (12 u ) \u2212K (12 t) 2 \u03c4 (12 t) ) . For a pair of gears, no meshing interference is the basic condition for normal operation. To ensure no meshing inter- ference, K (12) needs to remain non-negative, which means for an arbitrary , K (12) \u2265 0. Thus, from Eq. (57) , the necessary condition for meshing correctly can be derived as { K (12 t) \u2265 0 K (12 u ) \u2265 (\u03c4 (12 t) ) 2 K (12 t) . (59) If the relative normal curvature between (1) and (2) along n \u00d7 r (1) \u2032 is determined, from the geometry of a pair of tooth surfaces as shown in Fig. 6 , the contact ellipse can be derived. From Eq. (58) , the relative curvatures along the major and minor axes of any instantaneous contact ellipse, K ( a ) and K ( b ) , are derived as \u23a7 \u23aa \u23aa \u23a8 \u23aa \u23aa \u23a9 K (a ) = K (12 t) + K (12 u ) 2 \u2212 \u221a ( K (12 t) \u2212K (12 u ) 2 )2 + (\u03c4 (12 t) ) 2 K (b) = K (12 t) + K (12 u ) 2 + \u221a ( K (12 t) \u2212K (12 u ) 2 )2 + (\u03c4 (12 t) ) 2 . (60) The inclined angle of the major axis of an instantaneous contact ellipse from r (1) \u2032 about n can be derived as \u03b7 = \u2212\u03c0 4 \u2212 0 2 = \u22121 4 ( \u03c0 + arctan ( K (12 u ) \u2212 K (12 t) 2 \u03c4 (12 t) )) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003720_s0039-9140(96)02040-1-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003720_s0039-9140(96)02040-1-Figure2-1.png",
        "caption": "Fig. 2. Microcell with Nation membrane.",
        "texts": [
            " 1, (1) is the cylindrical graphite fiber working microelectrode, (2) is the reference/auxiliary stainless-steel needle, (3) is the sample, (4) is the hole for the micropipette for aspirating/dislodging a sample, (5) is the reference/auxiliary electrode contact and (6) is the working electrode contact. In the case of the anodic stripping voltammetry, it has been recommended [50] to preplate the reference/ auxiliary stainless-steel needle (2), with mercury. A two-electrode system of the catheter type has also been used in a microcell [60] but the electrodes were immersed in a supporting electrolyte in the outer compartment and this caused the undesirable dilution of a sample. The possibility of using a very small sample volume ( ~< 1/~1) in a microcell [44] (Fig. 2) has been reported with the working microelectrode placed in a Nation membrane tube. In Fig. 2 (1) is the working microelectrode (Au wire), (2) is the Nation membrane tube, (3) is the sample, (4) is the auxiliary electrode (Au wire), (5) is the reference electrode, (6) is the supporting electrolyte and (7) is a Teflon tube. Sample (4) and supporting electrolyte (6) enter the microcell by the capillary action. Since cations diffuse through the Nation membrane, only anions can be determined in the sample. Microcells with the working solid electrode in the form of a packed bed [15, 18, 19] have been developed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure51-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure51-1.png",
        "caption": "Fig. 51. Fluid Coupling Incorporating Planetary Gear",
        "texts": [
            " When the speed was increased to three times the original speed, the volume was reduced to about one-seventh, which meant that for a given filling pump the time for filling for manoeuvring and changing gear would be about one-seventh of that required in the first case. That feature was used in an arrangement now employed in certain cases for motor cars and for reversing gears and so on. 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 161 Last year he had used another method, Fig. 51, namely the division of the input power. One part was transmitted directly, and the remainder only was changed into hydrodynamic energy. This was effected by the well-known planetary gear. The effect was that a considerably higher slip might be chosen, with a relatively small loss, corresponding to about 2 to 2.5 per cent. The same coupling embodied a similar arrangement to that which the author had shown, with a movable scoop and the rotary ring vessel which was led round the whole gear. If it were desired to change from one gear ratio to another, that is to change from a converter to a coupling, it generally took about 10 seconds before the converter was filled, which was, of course, too long an interval for motor cars"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002900_j.isatra.2021.04.043-Figure23-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002900_j.isatra.2021.04.043-Figure23-1.png",
        "caption": "Fig. 23. This figure shows the experimental platform developed at the lab. The four rotors can tilt from approximately \u221230\u25e6 to 110\u25e6 . The aircraft is endowed ith the equipment depicted in Fig. 24.",
        "texts": [
            " [42] are: (a) an easy-to-implement control law in real convertible drones; (b) nonlinear aerodynamic model free of singularity is used instead of a linear one; (c) state convergence under external disturbances; and (d) a formal proof demonstrating global exponential stability. Besides, as was demonstrated, our results 5 We use the notation of V for the same variable. ( a f a 5 p a f c w 5 a it the transition region proposed by Naldi and Marconi, giving dditional support to our result. . Experiments This section implemented the presented controller in the real rototype CUAV depicted in Fig. 23. The conducted experiments re described in detail. To support our experimental results, we irst present a brief comparison with those works that have onducted experiments in a real convertible prototype. Let begin ith this. .1. Discussion of experiments presented in relevant works The control implementation in a real convertible prototype is n active area of research. According to state of the art, most works related to convertible aircraft do not provide real experiments. This can be seen in the following works published in the last years [40,44,57\u201360]",
            " The rest of the works shown in Table 6 only present flights n hovering. Having presented the previous discussion about the a ost important and relevant work of convertible aircraft that omprise experimental results, in the following, we pursue to xpose our obtained results in the CUAV prototype. .2. Experiments We present the experimental results showing how the transiion and back transition maneuvers are achieved stably according o the proposed approach. For that, we use the own designed nd built CUAV prototype depicted in Fig. 23. The transitions are t f s t 6 p c a r p t r t n l n t Fig. 24. Block diagram of the CUAV aircraft. The pilot gives the transition and back transition decisions. However, any desired signal can be triggered, such as a given state or even any sensor signal. performed autonomously, and it is activated by a signal provided by a decision-maker, in our case, a human pilot, although any other kind of command could generate it. Such a signal triggers when the transition starts and finishes according to function (66)",
            " The outputs of the saturation functions are given in Fig. 30. Fig. 31 shows the CUAV flying during the real-time flight tests. 6. Conclusion This work proposes a complete longitudinal aerodynamic mathematical model considering real aerodynamic terms for the Fig. 29. The control inputs\u2019 response for the complete flight task. Fig. 30. Saturation functions used in the transition control. Fig. 31. Transition mission: the CUAV during real-time flight tests. Convertible Unmanned Aerial Vehicle (CUAV) depicted in Fig. 23. Then, it is designed a smooth nonlinear controller to stabilize the CUAV in all flight envelopes, especially during the transition and back transition maneuver. We mathematically demonstrated that the CUAV\u2019s dynamics together with control (23) results to be GES. In particular, the control algorithm design takes into account the following characteristics: (a) The control algorithm takes the usually available information of conventional autopilots, such as IMU, pitot tube, and GPS. (b) The control implements saturation functions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002489_j.jwpe.2020.101671-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002489_j.jwpe.2020.101671-Figure1-1.png",
        "caption": "Fig. 1. The basic scheme of the biosensor architecture. (1) cathode layer, (2) membrane or separator, (3) anode layer, (4) open cover to fix the cathode layer, (5) anode chamber, (6) closed cover to fix the anode layer, (7) substrate inlet or outlet.",
        "texts": [
            " Carbon cloth (EC\u2212 CC1\u2212 060, Electrochem, Massachusetts, USA), was used as both anode and cathode, and a proton exchange membrane (Nafion N-324, DuPont, Texas, USA) as the separator in an ensemble arrangement (or as \u201csandwiched mode\u201d). Anode was pretreated by thermal oxidation before inoculation [3]. The original procedure was modified by lowering the temperature in half (318 \u25e6C). Nickel-Titanium (Ni-Ti, Ormco, California, USA) wires were threaded along the carbon cloth to be used as current collectors in both electrodes. Fig. 1 shows the basic scheme of the biosensor architecture. In a second experimental setup, the membrane was omitted and the cathode was pretreated by adding diffusion layers to avoid water leakage and minimize oxygen crossover to the anode [8,11]. A non-woven cloth was used as a separator to avoid contact between the electrodes. The MFCs were connected to a potentiostat-galvanostat (VSP Bio-Logic Science Instruments, EC-Lab ver. 10.12) to monitor the MFC voltage under closed-circuit operation. Voltage data were recorded automatically every 5 min"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001607_0954407020909663-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001607_0954407020909663-Figure4-1.png",
        "caption": "Figure 4. The hard points before optimization.",
        "texts": [],
        "surrounding_texts": [
            "Comparison of optimization variables and degree of understeer before and after optimization Comparison of optimization variables before and after optimization. Figure 6 shows that the height of the roll center of the front suspension after optimization can decline from the original 96.75 to 86.65mm. The roll steer coefficient is \u2202d=\u2202Fr. As can be seen from curves 1 and 2 in Figure 7, the coefficient before optimization is 0.131 and after optimization is 0.122. The roll steering coefficient decreased by 0.009, with little change. Comparison of equivalent cornering stiffness Cp before and after optimization. Table 9 presents the geometric suspension changes caused by the optimization of hard point and bushing stiffness have a direct influence on the equivalent cornering stiffness Cp of the suspension. Equivalent Cp of the front suspension changes from 82.69% to 75.63% after optimization, thereby falling into the reasonable range. Comparison of the degree of understeer before and after optimization. Table 10 shows that the tire cornering steer, roll steer, and lateral force steer are reduced to a certain degree via modification of the coordinates of the suspension hard points and bushing stiffness. Moreover, the degree of vehicle understeer is reduced, thus realizing the optimization of the excessive degree of vehicle understeer. Comparative analysis of vehicle minimum time handling and stability before and after optimization A simulation test on double-lane change is conducted using the ADAMS software. Based on the driver busy degree, rollover risk, side-slip risk indexes, and running and adhesion properties, single and comprehensive evaluation indexes of minimum time handling and stability are proposed in the previous literature.17,18 The formula is as follows ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s Figure 8 presents that the evaluation index of driver burden is reduced after optimization, and the degrees of driver busy and heaviness are mitigated. Figure 9 shows that after optimization, the evaluation index of rollover hazard is reduced. The roll angle is reduced Table 5. Sensitivity analysis of hard points. Hard point Roll center height (effect %) Roll steer coefficient (effect %) Hpl_tierod_inner x 1.97 2.14 Hpl_tierod_inner y 4.67 84.86 Hpl_tierod_inner z \u201396.33 \u201318.05 Hpl_tierod_outer x \u20138.43 13.68 Hpl_tierod_outer y \u20131.28 27.43 Hpl_tierod_outer z 91.2 \u201358.76 Hpl_lca_outer x 13.58 \u201329.78 Hpl_lca_outer y 4.63 \u201337.63 Hpl_lca_outer z \u201382.72 \u201342.56 Hpl_lca_front x 0.81 21.59 Hpl_lca_front y \u20133.86 6.47 Hpl_lca_front z 37.42 \u201335.42 Hpl_lca_rear x \u20131.66 22.14 Hpl_lca_rear y \u20136.38 17.85 Hpl_lca_rear z 58.07 33.83 Hpl_strut_lwr_mount x \u20132.96 \u201316.43 Hpl_strut_lwr_mount y 11.83 7.98 Hpl_strut_lwr_mount z 0.02 \u201370.32 Hpl_wheel_center x \u20130.53 52.60 Hpl_wheel_center y 5.73 \u20132.86 Hpl_wheel_center z 0.96 68.22 and rollover stability are improved. Figure 10 illustrates that the evaluation index of vehicle side-slip hazard is reduced. Hence, side-slip stability is further enhanced. Figure 11 shows that the evaluation index of dynamic is reduced. Vehicle speed and acceleration performance are improved, and vehicle ability on stable dynamic takeoff is reinforced. Figure 12 presents that the evaluation index of tire road holding is reduced. The tire camber angle is reduced, the adhesive rate is elevated, and the tire ground-grabbing ability is improved. Figure 13 illustrates that the responsiveness evaluation index is reduced, thereby improving vehicle acceleration response performance. Figure 14 shows that the comprehensive evaluation index of vehicle minimum time handling is reduced. Hence, minimum time handling and stability is improved."
        ]
    },
    {
        "image_filename": "designv11_14_0000300_bibe.2019.00122-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000300_bibe.2019.00122-Figure3-1.png",
        "caption": "Fig. 3. Structure of a inter-digital band-pass filter with optimization variables",
        "texts": [
            " Compared to parallel optimization methods such as PSO and GA which are non-gradient based, our proposed method is gradient based. For non-gradient methods the optimum solution is found usually at the expense of substantially more computational time compared to gradient based approach. The proposed parallel gradient method has higher speed of convergence over global optimization methods, better quality of the optimization solution and increased robustness over existing classical quasi-Newton method. Consider a standard inter-digital band-pass filter [22]as illustrated in Fig. 3. Assume equal spacing between each end of the resonator and the cavity wall. Coupling ratio between resonators are adjusted by tuning the values of spacing , spacing , and spacing between resonators. Each resonator is of length 43.18 mm, width 5 mm, thickness 0.5 mm, and the structure is enclosed in an cavity of height 10 mm. The design space vector for the example are chosen based on the sensitivity information. Fine model evaluation is performed by HFSS EM simulator using fast simulation feature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002669_s00366-020-01275-6-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002669_s00366-020-01275-6-Figure7-1.png",
        "caption": "Fig. 7 Folding mechanisms of a V-shaped unit (Type I) and b K-shaped unit (Type II)",
        "texts": [
            " Through careful observation of deformation pattern, two typical deformation modes are obtained during the whole deformation process of SHT tubes. These deformation modes may depend on the relative motion of adjacent cell walls. As for the two-panel (V-shaped) unit, the two walls move both to the same direction, one inward and the other outward. This belongs to the traditional in-extensional deformation mode (Type I). As for the four-panel (K-shaped) unit, all the four cell walls deform in the traditional extensional deformation mode (Type II), which can be seen in Fig.\u00a07b. Through analysis of deformation mechanism, it may be easily found that the plastic collapse of these SHT tubes satisfies the hypothesis of Simplified Super Folding Element (SSFE) theory. Fig. 6 Basic folding element: a in-extensional mode; b extensional mode 1 3 Figure\u00a08 gives the force\u2013displacement curves of SHT tubes with relative density \u2212 = 0.077 . It is depicted in Fig.\u00a08 that dynamic response curves of these triangular tubes show the common trend. All the dynamic response curves have three main phases: short elastic phase, plastic deformation phase, and finally densification phase [16]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001207_rpj-07-2018-0171-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001207_rpj-07-2018-0171-Figure1-1.png",
        "caption": "Figure 1 Applied load cases",
        "texts": [
            " During the design process, the race car did not exist, so the parameters for load transfer needed to be calculated or estimated (car weight, position of centre of gravity and magnitudes of G-forces). The weight of the car was calculated to be 250kg. Maximal longitudinal acceleration 2.1G and maximal transverse acceleration 3.1G were expected. These values were based on on-board sensors and diagnostics data from previous years. Both load cases of the right axle carrier of the formula are shown in working position in Figure 1. The axle carrier is connected to suspension rods (points marked A and B), whereas the steering rod is connected to point C. The colours are used to describe the applied constraints in the following way: The bolts (yellow in Figure 1) were preloaded. Constraints were applied on red parts, where the suspension rods (A and B in Figure 1) and the steering rod (C in Figure 1) aremounted: Constraint A: fixed in all directions, all rotations allowed; Constraint B: allowed movement in the Z direction, all rotations allowed; and Constraint C: allowed movement in theX and Z directions, all rotations allowed. Applied loads during the cornering load case: FG (1,200 N): weight, applied on bearings, in the centre of the wheel; and FC (3,730 N): cornering force, applied on bearings, in contact between wheel and road. Applied loads during the breaking load case: FG (1,200 N): weight, applied on bearings, in the centre of the wheel; FB (2,450 N): braking force, applied on bearings, in the centre of the wheel; and MB (575,000 Nm m): breaking momentum, applied on the brake calliper, free vector",
            "Model from previous design iteration was then used as a design space for next optimization. During a preliminary study on the previous version of the axle carrier [see the design space in Figure 2(a)], it was found out that more realistic results were obtained if the reaction forces were applied in the suspension pickup points and the constraints were applied on the bearings. Otherwise, the optimization algorithm did not connect the point for mounting the steering rod [Figure 2(b)], because constraints A and B (Figure 1) were sufficient for preventing the model from move and constraint C (Figure 1) was in that case redundant. If the forces were applied in the suspension pickup points, this phenomenon did not occur [Figure 2(c)]. Basically, the material was more functionally distributed near the forces than constraints. The finer material distribution was also obtained if the optimization was carried out in iterations and the material was removed gradually. Therefore three design stages were carried out using topology optimization until the weight was less than 500 g (Figure 5). After that, two minor design stages based on FEM evaluation were performed, and the shape of the axle carrier was optimized for additive manufacturing and defects in design were repaired",
            " The weight of the axle carrier in the first design stage was 875g. Additional restricted areas were added with respect to functionality, for example, the free space for the nut andwrench formounting [Figure 3(b)]. The second design stage was set on 30 per cent and theminimal wall thickness of 5.5mm.Theweightwas reduced to 596g. The specific behaviour of solidThinking Inspire (described earlier) was used in the third design stage. As the material is more functionally distributed near the forces, the loading of the axle carrier was applied as in FEM analysis (Figure 1), and the material distribution in the central part of the carrier was optimized. In this design stage, the supporting beam was added to prevent a flange fromdeformation [Figure 3(c)]. The basic shape of the axle carrier was then upgraded during three follow-up design stages. The component orientation on the building platform was chosen and the shape was additionally optimized with regard to additive manufacturing Topologically optimized axle carrier Ond rej Vaverka, Daniel Koutny and David Palousek Rapid Prototyping Journal Volume 25 \u00b7 Number 9 \u00b7 2019 \u00b7 1545\u20131551 and expected machining",
            " The dimensional accuracy of the axle carrier after additive manufacturing was checked by 3D scanning using the ATOS Triple Scan (GOMGmbH, Braunschweig, Germany). A special testing device was developed for the verification of real component deformation. It is described using Figure 8(a) in the configuration of the breaking load case. The main part of the testing device is a welded frame that allows the fixation of the axle carrier [blue in Figure 8(a)] at the suspension pickup points (A, B, C points in Figure 1). The loading is applied through the welded beam substituting real formula wheel [violet part in Figure 8(a)]. The beam is connected to the axle carrier similarly as the wheel using the bearings and brake Topologically optimized axle carrier Ond rej Vaverka, Daniel Koutny and David Palousek Rapid Prototyping Journal Volume 25 \u00b7 Number 9 \u00b7 2019 \u00b7 1545\u20131551 calliper. At the estimated wheel diameter, the loading forces (FG and FB) represented by threaded rods are applied. Loading forces are applied by tightening of the nuts in the direction of the arrows, as shown in Figure 8(a), whereas the loading forces are directlymonitored by strain gauges"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002107_jestpe.2020.3048091-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002107_jestpe.2020.3048091-Figure7-1.png",
        "caption": "Fig. 7. Experimental setup for PMSM torque control.",
        "texts": [
            " However, all the possible control inputs must be calculated step by step for backward recursion optimization in the dynamic programming, and transcendental equations must be solved when using Pontryagin\u2019s Maximum Principle. Those calculations are much more complicated than the definite three square root operations in RRCC. Therefore, the calculation amount in RRCC could be acceptable. In this section, a series of experiments is conducted to confirm the effectiveness of RRCC. The experimental setup is shown in Fig. 7. The electric dynamometer is used as the load of the PMSM under test. The parameters of the PMSM is shown in TABLE I. The tested PMSM torque control strategies are running at a frequency of 10 kHz on a piece of Infineon TriCore TC1782 which is the core processor of the control board. The PMSM is driven by a three-phase inverter using the SVPWM technique. The power semiconductor devices in the inverter are intelligent power modules, and the switching frequency is 10 kHz. The three-phase currents are measured by two LEM current transducers and converted by two onchip 12 bits A/D converters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001854_s00170-020-05828-3-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001854_s00170-020-05828-3-Figure1-1.png",
        "caption": "Fig. 1 Schematic illustration of PBF-LB/M process experiments using X-ray and thermal imaging in this study",
        "texts": [
            " The container was tapped 10 times to increase the packing density, which has been reported to be effective in reducing the balling effect [28]. The powder bed was flattened using a ruler. The mass of the container was measured without and with the powder, and the mass of the powder was calculated as the difference between these masses. The packing fraction \u03b1 was estimated to be 64% using Eq. (2). \u03b1 \u00bc M p \u03c1Ti6Al4VV s \u00f02\u00de where Mp is the mass of the powder, Vs is the volume of the sample holder, and \u03c1Ti6Al4V is the material density [29]. The specimen was placed inside a stainless steel vacuum chamber. Figure 1 shows a schematic illustration of the PBF-LB/M experiments conducted in this study using in situ X-ray and thermal imaging. Fused silica was selected as the window material, because it has a high transmission percentage of 90% across a wavelength range of 400\u20131200 nm, which permits the laser beam wavelength (1070 nm) and two infrared wavelengths (\u03bb1 = 800 nm and \u03bb2 = 975 nm) for the two-color pyrometer. The fused-silica window was located on the top of the vacuum chamber for passage of the laser beam, and it acted as an observationwindow for the two-color pyrometer",
            " (3) which was derived from Planck\u2019s equation: T \u00bc C2 1 \u03bb1 \u2212 1 \u03bb2 1 log R2 R1 \u22125log \u03bb1 \u03bb2 8>>< >: 9>>= >; \u00f03\u00de whereC2 = 1.438759 \u00d7 10\u22122 m \u00b7 K is the second radiation constant, \u03bbn (n = 1, 2) is the wavelength of the emitted light, and Rn is the radiation intensity at wavelength \u03bbn. The temperature was calibrated using a standard lamp, such that the temperature measurement error should be less than \u00b1 1% within the measurement range of 1300\u20132400 \u00b0C. The two-color pyrometer was placed approximately 570 mm in the perpendicular direction from the powder bed, as shown in Fig. 1. The size of the images was calculated to be 640 \u00d7 480 px with a pixel resolution of 13.3 \u03bcm at a distance of 570 mm. The entire inner width of the carbon holder was within the field of capturing area. The laser was placed on the left position of the thermal imaging. The variation of the spatial average temperature Tave with time was measured. During laser irradiation, X-ray imaging was used simultaneously with the two-color pyrometer for direct observation of the melt pool motion. The specimen was irradiated by an Xray source (MTT255: Shimadzu Corp"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002207_978-3-030-29131-0_5-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002207_978-3-030-29131-0_5-Figure11-1.png",
        "caption": "Fig. 11 Demonstration of the ACC on a highway scenario",
        "texts": [
            " a Recognition of walking activities and b Recognition of gait phases and events Wearable sensor for sit-to-stand activity\u2014The recognition of sit-to-stand (SiSt) and stand-to-sit (StSi) activities, has also been studied in laboratory environments usingwearable sensors. In the experiment shown in Fig. 10, the participant was asked to perform multiple repetitions of SiSt activity to collect acceleration data from one IMU. These data were processed by a probabilistic approach for recognition of sit, transition and stand states during the SiSt activity [47]. In addition,multiple segments that compose the transition state were recognised, which is important to achieve a robust and accurate control of assistive robots (Fig. 11). Other aspects that need to be considered in a real or outdoor environment could affect the recognition accuracy Fig. 10 Inertial measurement units employed for recognition of sit-to-stand and stand-to-sit activities Assistive Gait Wearable Robots \u2026 89 of SiSt. Some of these aspects are the delays in the pre-processing steps to smooth the signal, type of chair, optimal location of sensors and speed to move from sit to stand. Challengeswithwearablemeasurement systems\u2014Wearable sensors have been successfully used for recognition of human activities",
            " For this particular test case, a section of E40 highway scenario near Bertem in Belgium is virtualised using real world map data. The virtual scenario, sensor models and traffic simulation are all developed using Simcenter\u00ae Prescan. A dedicated button on the steering wheel is used for enabling/disabling the ACC functionality. In this scenario, the ego vehicle is cruising at approximately 100 km/h and approaching a leading vehicle. To avoid the collision, the ego vehicle slows down to the same speed as the leading vehicle as is shown in Fig. 11. After 41 s, the ego vehicle decides to overtake the leading vehicle. From this moment, the ACC increases the velocity of the ego to the previously defined setpoint. Autonomous Intersection Crossing The autonomous intersection crossing controller is demonstrated in the second scenario. The developed control algorithm is validated by having the ego vehicle approaching a cross-road together with one other road user. At a certain moment, their trajectories are overlapping and a collision might occur",
            " 10 Simulated contour error at the encoders and the TCP The simulation time for this contour is approximately 30 s. That is very efficient, although not yet real-time. In manufacturing technology, process stability and productivity tend to be the most vital properties for success. Coupling of a machine tool covering the whole mechatronic system with a production process model represents the top level of analysis. This is only possible with efficient, accurate, and low-order models. Kuffa [21] showed, by means of a MORe model of a test rig for high performance dry grinding as shown in Fig. 11, that this coupling is feasible. He divided the surface profile calculation in a machine-dependent part and a kinematic roughness part, which depends on the geometry of the grinding grains. Themachine dependent behaviourwas analysed bymeans of a simplified grinding process model. Good correspondence of the surface profile between measurements and simulation was achieved as shown in Fig. 12. An importantmeasure in precisionmachining is the volumetric accuracy of amachine tool. Volumetric accuracy is, according to Ibaraki and Knapp [22], represented by a map of position and orientation error vectors of the tool over the volume of interest"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001632_j.ijrefrig.2020.03.029-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001632_j.ijrefrig.2020.03.029-Figure3-1.png",
        "caption": "Fig. 3. Comparison of traditional and 360 \u00b0 circular injection port.",
        "texts": [
            " It is easy to suppress the blade vortex flow and improve he internal flow field of the impeller, which effectively reduces he impeller import and export losses ( Fig. 2 ). The impeller perfor- ance under different operating conditions is effectively ensured. .2. 360 \u00b0 circular refrigerant injection port The refrigerant injection port is designed as circular converence structure, which is arranged near the exit of backflow omponent. Comparison of traditional and 360 \u00b0 circular injection ort is showed in Fig. 3 . The traditional injection port delivers the ash steam into the internal flow of compressor directly, which ay produce strong disturbance on main flow channel and the on-uniform flow at the entrance of the second stage impeller. hrough the contraction design of gas supply channel, the static ressure is changed into dynamic pressure and intermediate gas is ccelerated to the same speed as main flow. Therefore, the main ow and the intermediate supply flow are in full contact on the ircumference and mixes in more fully and more evenly way"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001998_s00170-020-06296-5-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001998_s00170-020-06296-5-Figure2-1.png",
        "caption": "Fig. 2 Evaluation method for cylindricity and coaxiality of the stepped shaft: a cylindricity evaluation method; b coaxiality evaluation method. Where S1 and S2 are the measured elements of the upper and lower shafts, \u03d5T1 and \u03d5T2 are the minimum zone of the upper and lower shafts, L1 and L2 are the geometric axes of the upper and lower shafts, and S is the measured element of the stepped shaft, B is the reference element of the stepped shaft, \u03d5T is the minimum zone of the stepped shaft, and L is the reference axis of the stepped shaft",
        "texts": [
            " (4), then the measurement equation for stepped shaft with five systematic errors can be expressed as: \u03c1ij \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u0394rij \u00fe roj ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi cos2 \u03b7ij cos2 \u03b3\u00f0 \u00de \u00fe sin2 \u03b7ij vuut \u00fe r 0 B@ 1 CA 2 \u2212 d j1 \u00fe z j tan \u03d5\u00f0 \u00desin \u03b5\u00f0 \u00de \u00fe e j sin \u03b8ij\u2212\u03b1 j 2 vuuuut \u00fe e j cos \u03b8ij\u2212\u03b1 j \u2212r \u00fe z j tan \u03d5\u00f0 \u00decos \u03b5\u00f0 \u00de \u03b7ij \u00bc \u03b8ij \u00fe arcsin d j1 \u00fe z j tan \u03d5\u00f0 \u00desin \u03b5\u00f0 \u00de \u00fe e j sin \u03b8ij\u2212\u03b1 j \u0394rij \u00fe roj ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi cos2 \u03c6ij\u2212\u03b2 cos2 \u03b3\u00f0 \u00de \u00fe sin2 \u03c6ij\u2212\u03b2 vuut \u00fe r 0 BBBBBB@ 1 CCCCCCA \u2212\u03b2 8>>>>>>>>>< >>>>>>>>>>: \u00f010\u00de As shown in Fig. 2a, the minimum zone method is used to evaluate the cylindricity. According to the structural characteristics of the stepped shaft, the paper designs a coaxiality measurement scheme as shown in Fig. 2b. The lower section of cylinder is selected as the reference element, whose axis is used as the reference axis of the stepped shaft. The upper shaft section is the measured element to calculate the coaxiality of the stepped shaft. As shown in Fig. 1, if there is no measurement error, the sample angle \u03b8ij is uniform, but in the actual measurement, systematic errors, such as eccentricity and probe offset, existed, resulting in sample angle deviation and causing the real sample angle \u03c6ij to be non-uniform distribution, and Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001905_j.mechmachtheory.2020.104095-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001905_j.mechmachtheory.2020.104095-Figure7-1.png",
        "caption": "Fig 7. Coordinates of the hybrid compliant mechanism.(a)3-D view (b) top view.",
        "texts": [
            " 3) The rotation angle of the central chain is considered as a small angle. 4) To facilitate dynamic modeling, in the following work, we regard the mass block as a part of the moving platform. 2.2. Kinematic model of kinematic chains Section 2.2 consists of 3 steps. In step 1, we define coordinates and symbols used in this section. Then, in step 2, we set kinematic chain 1 as an example to introduce the PR PRBM of the kinematic chain. Finally, in step 3, we give the coordinates of the mass center of rigid links in PR PRBM. Step 1 . As shown in Fig. 7 , five coordinates are established in the hybrid compliant mechanism. The global coordinate { O B }- O B x B y B z B is fixed at the base. The moving coordinate { O P }- O P x P y P z P is connected to the moving platform. The symbols used in this paper are defined as follows: P i , i = 1,2,3 represent the connection points between the kinematic chain and the moving platform; B i , i = 1,2,3 denote the kinematic installation position on the base; R is the transformation matrix from { O P } to { O B } shown in Eq",
            " To facilitate showing the constraint relationship on plane x B O B z B , we give an axonometric drawing in Fig. 10 (b). In Fig. 10 (a), triangle aO p \u2019b is similar to triangle bP 3 c . We get \u03b11 = aO p \u2019b . Line bc is parallel to axis z B . That makes \u03b81 x = aO p \u2019b . Then, considering that \u03b1 = aO p \u2019b , we can get the relationship shown in Eq. (21) . \u03b81 x = \u03b11 = \u03b1 (21) Next, considering that triangle P 2 dP 3 is a right triangle, we can get the relationship shown in Eq. (22) . sin \u03b11 = P 2 d P 2 P 3 = z p2 \u2212 z p3 \u221a 3 L (22) where, the meaning of L is shown in Fig. 7 (b). In Fig. 10 (b), P c is the center of P 0 P 2 . That makes z pc = 0.5( z p 2 + z p 3 ). We can find that triangle aO p \u2019b is similar to triangle bP 3 c , line bc is parallel to axis z B and \u03b1 = aO p \u2019 b . Then, we can get Eq. (23) \u03b81 y = \u03b21 = \u03b2 (23) Triangle P 1 dP c is a right triangle, that means we can get Eq. (24) . sin \u03b21 = P 1 d P 1 P c = z p1 \u2212 0 . 5 ( z p2 + z p3 ) 1 . 5 L (24) Considering that the deformation of the central chain (about 30 \u03bcm) and the cantilever beam (about 0.03 \u03bcm) is not significant, the value of \u03b81 x, \u03b81 y , and \u03b82 can be regarded as insignificant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure2-1.png",
        "caption": "Fig. 2. D\u2013H parameters of the 6R mechanism.",
        "texts": [
            " From Fig. 1, we obtain |AB| = 2|CA| sin(\u2220ACB/2) = 2|DA| sin(\u2220ADB/2). (1) It is noted that \u2220ACB=\u2220ACB1+\u2220B1CB=\u2220A1CA+\u2220ACB1=\u2220A1CB1, |CA1|=|CB1|, \u2220ADB=\u2220ADA2+\u2220A2DB=\u2220A2DB+\u2220BDB2= \u2220A2DB2, |DA2|=|DB2|, |CA| = \u221a|CA1|2 + r2 and |DA| = \u221a|DA2|2 + r2. Eq. (1) then becomes \u221a |CA1|2 + r2 sin(\u2220A1CB1/2) = \u221a |DA2|2 + r2 sin(\u2220A2DB2/2). (2) Different variations of D\u2013H notations are used in the literature [8,12,14,31\u201333]. The coordinate frames are attached to the links and the link parameters are defined as in Ref. [33] (Fig. 2). Zi-axis is along the axis of joint i. Xi-axis is along the common perpendicular between Zi\u22121- and Zi-axes. Oi is the intersection of Xi- and Zi-axes. Yi-axis is defined by Xi- and Zi-axes through the right handed rule. The link parameters of link i are: di (the distance between Xi- and Xi+1-axes measured from Xi-axis to Xi+1-axis along Zi-axis), ai (the twist angle between Zi- and Zi+1-axes measured from Zi-axis to Zi+1-axis about Xi+1-axis), and li (the distance between Zi- and Zi+1-axes measured from Zi-axis to Zi+1-axis along Xi+1-axis)",
            " Since the approaches to the kinematic analysis of 6R mechanisms have been well documented in the literature (see [6,7,12,14,20,31,32,34,35] for example), detailed derivation will be omitted in this paper. The method for kinematic analysis in Ref. [8] is not used in this paper since it requires one to first determine the link parameters of two 5R Goldberg mechanisms and then perform the kinematic analysis of these two 5R Goldberg mechanisms. The link parameters of an example mechanism \u2014 Mechanism I (Fig. 2) \u2014 are given as follows: l1 = l2 = l4 = l5 = 120, l3 = 0, l6 = 0, a1 = a2 = a4 = a5 = p/2, a3 = p/3, a6 = p/2, d2 = d5 = 0, and d1 = \u2212d6 = 50. Using Eq. (3), we obtain the link parameters d4 and d3 as d4 = \u2212d3 = 10 \u221a 194. The input\u2013output equation of Mechanism I is derived as \u2212169Sh1Sh2 \u221a 194 + 2028Sh1Ch2 + 119 \u221a 194Ch2 + 120Sh2 \u221a 194 + 2028Sh1 \u2212 1440Ch2 + 1428Sh2 \u2212 845 = 0. (4) For a given set of h1 and h2, one can determine hi (i = 3, 4, \u00b7 \u00b7 \u00b7 6) by calculating ti = cot(hi/2) using the following equations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000277_1350650119896455-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000277_1350650119896455-Figure1-1.png",
        "caption": "Figure 1. Photos of the stator and the rotor before testing. (a) Stator; (b) rotor.",
        "texts": [
            " The cryogenic experimental results are very helpful in guiding the design of high-performance shaft seals for the high-speed LH2 and LO2 turbopumps. Test object and experimental parameters For finishing the cryogenic tests of the mechanical face seal, a series of experiments theoretical studies as shown in Zhang et al.22,23 have been carried out. With the theoretical results, the optimizations on the structure parameters of the tested seals under the cryogenic conditions have been completed, and the tested seals described in this paper were designed. Figure 1 shows photos of the stator and the rotor before testing. The stator is composed of a foundation bed and a copper-graphite ring. The material of the foundation bed is aluminium bronze. There is a ring channel with a width of 10mm and a depth of 1\u20132mm on the top of the foundation bed, and a 3mm-thick copper-graphite ring is installed in the channel by brazing the bed and the graphite ring; 9Cr18 stainless steel is used as the rotor material. Low-temperature cryogenic treatment of the rotor was carried out before testing. As shown in Figure 1, the structure of the rotor has external spiral grooves and internal herringbone-shaped grooves. Based on the results of previous experiments,22 the closing force of the seal in this paper is defined as being less than 300N. In this study, the flexible support of the stator is provided by 6 springs, and the compressed length and stiffness of the spring are 8.8mm and 3.04N/mm, respectively. Hence, the total spring force Fz is 160N. Table 1 presents the main running and geometric parameters of the tested mechanical face seal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000281_iecon.2019.8927827-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000281_iecon.2019.8927827-Figure16-1.png",
        "caption": "Fig. 16 \u2013 Bonded magnets used in axial flux machine: rotor (a) and stator (b)",
        "texts": [
            " The output voltages of the machines equipped with the two considered magnets are identical (Fig. 15). Where compact geometries and weight reduction become a challenging constraint, the powder material technology can be considered, both for the stator soft magnetic structures (in general, Soft Magnetic Composite are adopted), and for the permanent magnets, allowing different freedom degrees to the machine designers. The bonded magnets were efficiently used in a 1.5 kW synchronous axial flux machine (Fig. 16), resulting the best choice for the prototype production. The mass production also benefits from the simplicity of the whole process, that only requires the implementation of a suitable mould. The torque-current characteristic has been reported in Fig. 17, the measurement trend has good matching with the simulation and analytical methods. Another prototype of an axial flux machine is shown in Fig. 18. The very compact rotor and stator shape gave the best results, leading to an efficiency of 0.81 for a 150 W rated machine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003885_7.624348-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003885_7.624348-Figure1-1.png",
        "caption": "Fig. 1. Flexible SPICE structure.",
        "texts": [
            " The hexagonal base referred to as the bulkhead forms the support for the entire structure and is 6.19 m in diameter. The primary mirror (PM) assembly is mounted on top of the bulkhead. Three legs (tripod) connect the bulkhead to the secondary mirror (SM) assembly. The SM Assembly is 1.32 m in diameter. The overall height of the structure is 8.14 m. Maintaining the alignment of the SM Assembly and the bulkhead is the primary concern of this research. An exaggerated example of the SPICE structure exhibiting a misalignment due to its flexible bending modes is illustrated in Fig. 1. Note that (for the purposes of this research) the alignment is not altered by a pure torsion force about the LOS axis. Actuators provide the control force necessary to quell the structural vibrations. The specific actuator utilized is referred to as a proof mass actuator (PMA). The PMA uses a proof mass that is electromagnetically moved to counteract the bending motion of the structure at the location of the PMA. Eighteen PMAs are mounted on the structure. There are 6 PMAs located such that there is one on the vertical spar at each of the hexagonal corners of the bulkhead, each pointing in the Z direction",
            " Since there were no changes made between the truth model and filter models for dynamics driving noise or the measurement noise inputs to represent actual disturbances in the real world system, the primary method of tuning Qf and Rf (to address order reduction of the filter design model) was by adjusting each of the respective values by a scalar multiplier. This procedure worked well, although there is a noticeable difference in the tuning of the X and Y axis, respectively, due to the geometrical differences in the two axes on the actual SPICE tripod structure (reference Fig. 1). For controller tuning [9, 11], only the scalar weight affecting the state weighting matrix was adjusted, since the value for the control weighting matrix was assumed to be at its limiting admissible value [1]. Tuning the controller was a simple matter of adjusting the scalar state weighting value until the point of closed-loop system instability was reached (determined visually by a divergence in the error estimation plots or the LOS plots). This point implies the \u201ctightest\u201d possible state values will be obtained with slightly smaller state weightings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure5-1.png",
        "caption": "Fig. 5. Transition pose of 3-R(RRR)R+R HAM in azimuth motion: (a) Transition pose I; (b) transition pose II.",
        "texts": [
            " 3, the expected pose of pitch motion is selected as the initial pose of azimuth motion and Fig. 4 is set as the desired pose of azimuth motion. In the course of azimuth motion, the HAM changes from the initial pose to the expected pose and the azimuth changes from \u03c2 to \u03c21. To keep the direction of vector T V always along the intersection of plane OCT and plane Tx T Ty , the rotation angle of the polarization mechanism is \u03b86. Two sets of transition poses can be used to realize the expected pose of the HAM from the initial pose of the azimuth motion to the expected pose, as shown in Fig. 5. https://doi.org/10.1017/S0263574720001290 Downloaded from https://www.cambridge.org/core. University of Toledo, on 03 Jun 2021 at 19:25:40, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. The motion process is as follows: (a) the HAM keeps the azimuth angle unchanged from the initial position and attitude, only carries out pitch motion, and the end coordinate system T \u2212 Tx Ty Tz does not rotate relative to the moving coordinate system C \u2212 CxCyCz until it reaches the transition pose I as shown in Fig. 5(a), (b) the axis revolute angle of the end coordinate system T \u2212 Tx Ty Tz around the moving coordinate system C \u2212 CxCyCz is \u03b86 = \u03c21 \u2212 \u03c2 . The pose of the antenna reflector reaches the transition pose II as shown in Fig. 5(b), (c) keeps the azimuth angle of the moving platform unchanged, and continues to pitch to move upward until the expected pose of the azimuthal movement is achieved. The HAM moves from the initial pose to the transition pose I and transition pose II to the expected pose, and the HAM does not rotate when it is pitching; when the HAM moves from the transition pose I to transition pose II, the polarization mechanism rotates. The 3-R(RRR)R+R HAM can solve the problem of the phase change of the reflector caused by the accompanying motion by introducing a single degree of freedom polarization mechanism and meet the working requirements of non-circular polarization antenna"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001526_s00170-020-04987-7-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001526_s00170-020-04987-7-Figure8-1.png",
        "caption": "Fig. 8 Void/material border position error inherent to the voxel size",
        "texts": [
            " The volume was then reconstructed using the CT Pro 3D software (Nikon, Brighton, MI, USA). Due to its height, the part was scanned in two acquisition windows and assembled after the reconstruction to get the entire part. Note that despite the segmentation quality and the software accuracy, there are some limitations in the default detection, inherent to the voxel size. Since a voxel has only one intensity, the 52\u03bcm\u00d7 52\u03bcm\u00d7 52\u03bcmvolume contains only one piece of information. When the voxel is at a void/material border, there is an error induced by its size (Fig. 8). An STL was reconstructed using the segmented part and compared to the CAD, in order to extract the overall 3D profile deviation detected by the \u03bc-CT. Figure 9 presents a setup used to scan the case study part, along with a scan 3D rendering. The obtained scan is presented as a series of voxels having different grayscales. A threshold is applied in the Dragonfly V2 software (Object Research Systems, Montreal, Canada) environment to separate the region of interest (ROI) which is the part, from the void"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001312_tie.2019.2952780-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001312_tie.2019.2952780-Figure3-1.png",
        "caption": "Fig. 3. Schematic diagram of the single integrated electrically driven spindle and clamping Actuator: (a) Headstock housing different parts, (b) Electromechanical Actuator in isomeric view, (c) XY plane 2D representation of the electromechanical Actuators.",
        "texts": [
            " 2) The electromechanical drive mechanism must be of a size that can be installed on the headstock of the existing hydraulically driven system. 3) Simultaneous powering of the clamping and the spindle system by the driving mechanism should be avoided to reduce the overall energy consumption. Once the clamping operation is over, the power to it should be disconnected and it should be capable of maintaining the clamping force. By ensuring this, power should only be delivered to the spindle. The headstock containing the proposed electromechanical actuator is shown in Fig. 3(a). Fig. 3(b) and 3(c) show the isomeric and the YZ plane view of the electromechanical actuator. The electromechanical actuator consists of a clutch, an epicyclic gear and a screw mechanism which converts the rotational motion into a linear motion and performs the clamping and the constant speed spindle rotation integrally with the help of the single drive motor attached to it. The detailed operation of the electromechanical actuator is further explained in the following subsections. Fig. 4 highlights the clutch and the spindle mechanism of the proposed electromechanical actuator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002362_tie.2020.3000088-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002362_tie.2020.3000088-Figure15-1.png",
        "caption": "Fig. 15. Bench of PM machine saliency-based sensorless drive",
        "texts": [
            " A delay time T1, with 1~10 microsecond is added to avoid any sampling errors during the controller initialization. T2 is the stop time for accumulation. Dependent on different machine saturation conditions, the overall accumulation process might spend around hundreds millisecond. Detailed results will be shown in the following experimental section. IV. EXPERIMENTAL RESULTS A 2kW 8-pole interior PM machine with 3000rpm rated speed is tested to evaluate the proposed PWM voltage injection for the saliency-based sensorless drive using lowside shunt current sensing. Fig. 15 shows the machine drive test bench. The test machine is coupled to a load machine for the torque operation. Key machine characteristics are listed in Table III. Regarding to the current sensing hardware, a 2m\u03a9 shunt resistor is selected to measure the phase current at a full scale of 8.5A. Three amplifiers with 50MHz bandwidth are selected. The detailed current sensing specification is shown in Table IV. All sensorless drive algorithms are implemented in a 32-bit microcontroller, TI-TMS320F28069"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure1-1.png",
        "caption": "Fig. 1. Pump-Controlled Scoop Tube Coupling",
        "texts": [
            " VARIABLE-FILLING FLUID COUPLINGS Variable-filling couplings are most commonly used for speed regulation in connexion with constant-speed driving motors, control of the quantity of liquid in the working circuit being effected by means of a -~ * Managing Director, Hydraulic Coupling and Engineering Company, Ltd. [I .MECH .E .] 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from scoop tube which maintains a constant circulation through an external circuit of liquid drawn off from the working chamber. The external circuit provides a convenient means by which liquid can be added or removed by a reversible pump, to increase or reduce the speed as desired. Fig. 1 shows such a design, known as a\u201cpump-controlled scoop tube coupling,\u201d having a cooler in the external circuit. 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID .COUPLINGS 85 The working circuit formed by the curved passages of the impeller and runner of the scoop tube couplings has for many years been of the general form originated by Professor Fottinger, and in its most efficient proportions is adopted in the \u201cVulcan\u201d coupling. The object of the core guide ring, around which the vortex ring of liquid circulates, is to ensure the highest efficiency in consequence of an orderly flow of the liquid with the lowest frictional and eddy losses when the coupling is full",
            " An improved scoop tube with thin walls and a taper tip was successfully introduced to deal with the required increase in the circulating pressure, and its capacity was proved by tests upon experimental couplings to be easily sufficient for the new requirements. After installing the new scoop tube in a number of couplings, a curious inconsistency was observed in their operation, although they were identical in design. Some of the couplings tended to lose speed slowly when driving for long periods boiler fans working at constant 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 92 PROBLEMS OF FLUID COUPLINGS load. At first, tightness of the suction and delivery foot valves in the reservoir tank (X in Fig. 1) was suspected. In a number of cases leakage was proved to be due to distortion of the ball valve seatings, 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 93 which had been pressure-tight under test, as a result of the effects of time and temperature on the cast iron valve body. The foot valve design was altered and tests made with various materials for the seating, the best results being obtained with a high-grade whitemetal in a mild steel housing",
            " 11, however, the scoop tube is situated in a 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from region where no liquid would be picked up by such a scoop under the ordinary fan drive conditions described, and hence it does not appear possible to give any useful range of speed regulation with this design. The reliability of a main power transmission is of course determined by the reliability of any auxiliaries essential to its operation. In the case of the pump-controlled scoop tube coupling shown by Fig. 1, a very high standard has been achieved by using a simple reversible gear pump which is normally idle, and is Elimination of Auxiliaries. operated only during the moments when the speed is being adjusted. I t is, however, in the case of automatically controlled drives requiring a continuously running servo-pump and piston valve for filling or emptythat the question of dependence upon auxiliaries is a reasonable point of criticism. An improvement in this direction is the displacement cylinder control evolved in conjunction with the Hagan automatic regulator, as illustrated by Fig",
            " The degree of braking is under the direct control of the driver and is smooth and progressive without risk of skidding the wheels, but the mechanical brakes would of course have to be relied on to bring the vehicle to rest and hold it stationary. Mechanical Troubles. The record of mechanical troubles is not a long one in the case of the pump-controlled scoop-tube coupling, as might be expected in view of its simple mechanical design. In one boiler house installation a number of failures occurred with the earlier design having a spigot ball bearing supporting the runner stub shaft in the boss of the impeller. This bearing is mounted in a slightly resilient steel housing (called by the erectors a \u201ctin hat\u201d), Y in Fig. 1, p. 84, which is effective in compensating for small angular errors in alignment. The * By J. M. Voith Maschinenfabrik, Heidenheim. 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 107 trouble appears to have been due to misalignment, accentuated by the compact layout of the motor and combination bedplate supporting the two adjacent bearings of the fan shaft. Curiously enough, a similar degree of misalignment was -present in the induced draught fan couplings in the same boiler house, but since these were of the double-inlet type with the bearings spaced perhaps 15 feet apart, the flexibility of the long fan shaft and the resilience of the \u201ctin hat\u201d were sufficient to keep the load on the spigot bearing within tolerable working limits",
            " The torque was limited to a predetermined value, despite the many turns and gradients included in the full circuit of the endless chain, which might be a mile or more long. In 1933, he was responsible for the construction of a conveyer 800 feet long between the end shafts, for taking bags of sugar from the warehouse of a London refinery to the wharves. The conditions of working were such that the conveyer had to be frequently started and stopped when fully loaded with sugar, causing repeated shocks to the transmission. It was decided to fit the existing 16 h.p. motor, with a fluid coupling of the scoop tube type, as shown in Fig. 1, p. 84, with pump * PROCEEDINGS, 1935, vol. 130, p. 75. 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from and oil tank but without the cooling device. It was installed without scrapping any existing gearing, it eliminated all starting shocks. from the conveyer chains and mechanism, and it furnished a simple method of adjusting the speed of the conveyer to suit the conditions of working. I t therefore became easy to hold back the bags of sugar when they were coming along too quickly to be handled, and also to accelerate the supply when required"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000461_etep.2171-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000461_etep.2171-Figure2-1.png",
        "caption": "Figure 2. Rotor equivalent circuit.",
        "texts": [
            " Td\u00bd \u00bc \u00fe1 1 0 0 1 1 1 0 \u00fe1 2 64 3 75 (3) The matrix of resistances is expressed Rsd\u00bd \u00bc ras rbs 0 0 rbs rcs ras 0 rcs 2 64 3 75 (4) ras, rbs, and rcs are the resistances of stator windings. The stator flux is defined by the following: \u03a6sd\u00bd \u00bc Td\u00bd \u03a6s\u00bd (5) \u03a6s\u00bd \u00bc \u03a6as \u03a6bs \u03a6cs\u00bd T The rotor cage is composed of n bars and the end ring circuit. It is modeled by an equivalent circuit containing n magnetically coupled circuits. Each rotor loop consists of two adjacent bars and the two portions of the end ring connect them as shown in Figure 2. Because the stator and rotor currents contain high frequency components, the skin effect (deep bar effect) causes the equivalent rotor circuit resistance and leakage inductance to vary with the harmonic frequencies [28,29]. Applying the superposition principle, the equivalent rotor resistance can be defined by the fundamental and harmonic resistances, which are placed in parallel. Each harmonic resistance corresponds to the harmonic current of the same order. However, the harmonic currents are very small in comparison with the fundamental current"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001905_j.mechmachtheory.2020.104095-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001905_j.mechmachtheory.2020.104095-Figure9-1.png",
        "caption": "Fig. 9. 1R PRBM of central chain and cantilever beam: (a) Plane x B \u2019O B \u2019 z B \u2019 (b) Top view.",
        "texts": [
            " In step 1, we introduce the 1R PRBM of the central chain and the cantilever beam. Then, in step 2, we give the kinematic model of the central chain and the cantilever beam. Step 1. Based on the analysis in Section 2.1 , the pose of the central chain and the cantilever beam can be described by rotation angles. That means the 1R PRBM could be utilized to describe the deformation of the central chain and cantilever beam. Considering that the central chain could be bent in any direction, the deflection coordinate { O B \u2019 }- O B \u2019x B \u2019y B \u2019z B \u2019 shown in Fig. 9 is set up to describe the deflection of the central chain. Fig. 9 also show the 1R PRBM of the central chain and the cantilever beam. To facilitate modeling, we consider the deflection direction of the cantilever beam is consistent with the central chain. Considering that the young\u2019s modulus of the moving platform (about 7.8GPa) is much larger than the central chain (about 0.0078GPa), the moving platform could be represented by lumped mass. According to [39] , the 1R PRBM is composed of two rigid links and a revolute (R) pair with a torsion spring. The length of each rigid link in the 1R PRBM is presented as \u03b3 ic l ( i = 0,1). k ic ( i = 1,2) is the torsion spring stiffness. These factors and coefficients of 1R PRBM are obtained in [39] and can be presented as follows. \u03b30 c = 0 . 15 \u03b31 c = 0 . 85 k c = 2 . 65 k ic = \u03b31 c k c EI l i = 1 , 2 For central chain, E = 0.0078Gpa I = 490.87mm 4 l = 40mm. For cantilever beam, E = 209 Gpa I = 0.049 mm 4 l = 30mm. Other variables in Fig. 9 are defined as follows. l ic ( i = 1,2,3,4) is the length of rigid links in Fig. 9 . l Mc is the height of the moving platform. m ic ( i = 1,2,3,4) means mass and J ic ( i = 1,2,3,4) means mass moments of inertia of rigid links, respectively. m M is the mass of the moving platform. J Mc is the mass moment of inertia of the moving platform. The subscript M means moving platform. The subscript c indicates that the symbol is related to the central chain or the moving platform or the cantilever beam. \u03b81 and \u03b82 are rotation angles of the revolute pair in 1R PRBM. The symbols in Fig. 9 are presented in Table 3 . Step 2. Then, according to Fig. 9 (a), the mass center of rigid links I, II, III, IV and the moving platform can be expressed as below: { B \u2032 x I = 0 B \u2032 z I = 1 2 l 1 c (12) Table 3 Symbols in Fig. 9 . Symbol m 1 c m 2 c m Mc m 3 c m 4 c Value 0.42g 2.56g 5.76g 0.028g 0.156g Symbol J 1 c J 2 c J Mc J 3 c J 4 c Value 4.185g \u00b7 mm 2 262.976g \u00b7 mm 2 2244.3g \u00b7 mm 2 0.048g \u00b7 mm 2 8.475g \u00b7 mm 2 Symbol l 1 c l 2 c l Mc l 3 c l 4 c Value 6mm 34mm 18mm 4.5mm 25.5mm { B \u2032 x II = 0 . 5 l 2 c sin \u03b81 B \u2032 z II = l 1 c + 0 . 5 l 2 c cos \u03b81 (13){ B \u2032 x M = ( l 2 c + 0 . 5 l Mc ) sin \u03b81 B \u2032 z M = l 1 c + ( l 2 c + 0 . 5 l Mc ) cos \u03b81 (14){ B \u2032 x III = ( l 2 c + l Mc + 0 . 5 l 3 c ) sin \u03b81 B \u2032 z III = l 1 c + ( l 2 c + l Mc + 0 . 5 l 3 c ) cos \u03b81 (15){ B \u2032 x IV = ( l 2 c + l Mc + l 3 c ) sin \u03b81 + 0 . 5 l 4 c sin ( \u03b81 + \u03b82 ) B \u2032 z IV = ( l 2 c + l Mc + l 3 c ) cos \u03b81 + 0 . 5 l 4 c cos ( \u03b81 + \u03b82 ) (16) Point P 0 and O p \u2019 can be expressed, respectively: { B \u2032 x P 0 = ( l 2 c + l Mc + l 3 c ) sin \u03b81 + l 4 c sin ( \u03b81 + \u03b82 ) B \u2032 z P 0 = ( l 2 c + l Mc + l 3 c ) cos \u03b81 + l 4 c cos ( \u03b81 + \u03b82 ) (17){ B \u2032 x O \u2032 P = L c sin \u03b81 B \u2032 z O \u2032 P = l 1 c + L c cos \u03b81 (18) where, L c = l 2 c + 0 . 45 l Mc . Then, according to Fig. 9 , we can get Eqs. (19) and (20) B \u2032 x 2 O \u2032 P = L 2 c sin 2 \u03b81 = X 2 + Y 2 (19){ Y = L c sin \u03b81 x X = L c sin \u03b81 y (20) 2.4. Constraint relationship between central chain and kinematic chains In this section, we give the constraint relationship between the pose of the moving platform and the generalized coordinates of kinematic chains. In Fig. 10 , we project the deformation of the central chain and the cantilever beam, the pose of moving platform to plane y B O B z B and plane x B O B z B , respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002904_tmech.2021.3074800-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002904_tmech.2021.3074800-Figure3-1.png",
        "caption": "Fig. 3. The thin-film FSRs underneath the handguards.",
        "texts": [
            " As the user performs the STS transfer, the linear actuator extends and the support subsystem rises. The assistive force/torque is transferred to the user through the harness and the support plate. In the assistive walker, the motion of the linear actuator should be synchronized with the human behavior in accordance with the user-walker interaction. For appropriate interaction measurement during the STS transfer, Jun [3] used a 6-axis force/torque sensor to measure the compression force. Similarly, we installed a thin-film force sensing resistor (FSR) underneath each handguard, as shown in Fig. 3, to detect the compression force between the arm and the support plate. The user\u2019s intention of standing up is measured through the magnitude of the compression force. To measure the user\u2019s posture, an integrated wireless wearable IMU sensor system proposed in [36] is used in this paper. The sensor nodes of the system are worn at the trunk, thigh and shank levels of the subject\u2019s body as shown in Fig. 4. These nodes can measure the attitude angles of body segments with a frequency of 20Hz. The attitude angles consist of yaw angle \u03c8, roll angle \u03c6 and pitch angle \u03b8"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002419_j.ijepes.2020.106382-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002419_j.ijepes.2020.106382-Figure2-1.png",
        "caption": "Fig. 2. Two-mass drive train iconic diagram.",
        "texts": [
            " According to (41), in the pu system, effort, flow, resistance, inertia, capacitance, generalized displacement, and generalized momentum on either side of a TF element have equal magnitudes; in other words, these per-unit elements are related by a virtual TF element that has a modulus equal to one =m( 1)i , whereas maintaining the junction structure as that of the original bond graph. Consequently, pu parameter influence on state variables is readily available. This procedure may be applied to bond graphs that have a Weighted Junction Structure with no modulated transformer (MTF) elements [2]. It is assumed that if bond loops exist in the BG, then these bond loops are simple meshes [2]. Consider the drive train iconic diagram depicted in Fig. 2, where g and t are the rotary mass angular speeds of the generator and turbine respectively in rad/s, g and t are the instantaneous generator and turbine angular positions in rad, is the spring angle of torsion also in Trad, K12 is the torsionally elastic torque component of the shaft in N m which is related to by = K T n,K p12 12 is the number of poles of the generator, and N N/1 0 is the gear ratio. The associated two-mass drive train model is often found to be sufficiently accurate to be used in transient response studies [5\u20137]. Nomenclature and values are given in Table 1 for a wind turbine with a 87.965 gear ratio and a 4-pole generator. Using the BG methodology for Fig. 2 and employing the classical force-voltage analogy [2], the bond graph shown in Fig. 3a was constructed [2,3], to which integral causality has been assigned as the preferred causality according to the sequential causality assignment procedure [2, p. 103]. Applying to Fig. 3a the procedure for manual derivation of equations from causal BG (see [2,3]), the following state equation is derived = + + t T K T T T d d 0 0 0 0 0 g t K D D J D n J J D n J D J D n J n J K n g t K n J J e W 1 1 12 1 1 12 1 12 1 12 1 1 1 12 1 2 2 2 12 1 2 2 1 2 12 1 12 0 1 2 (42) Fig",
            " Likewise, according to (74) and (79), the pu shaft stiffness may or may not be multiplied by b. This is a possible source of confusion in the use of the pu system when applied to the rotational mechanics energy domain, i.e., in the pu system, K and K\u0304 are not, and in fact cannot both be regarded as being expressed in pu excepting for the exceptional case in which the angular displacement base value is taken as b b (compare e.g., [8,13]). Real quantities of the wind turbine two-mass drive train shown in Fig. 2 correspond to the WindPACT 1.5 MW turbine [12] and are shown in Table 1. Parameters D1 and D2 have been proposed for the sake of example. Selecting = \u00d7P 1. 5 10 Wb 6 as the power base, = =q 1 radb b (0) (0) , and = =f 376. 9911 rad/sb b (0) (0) , the effort base value for stage (0) becomes = = = = \u00d7e T P f P 3. 9789 10 N mb b b b b b (0) (0) (0) (0) 3 (80) By utilizing (43)\u2013(51), the following base values can be calculated = = = \u00d7 = \u00d7 = \u00d7 = \u00d7 = \u00d7 = = \u00d7 = \u00d7 = = \u00d7 T T K D K K D D 188. 4956 rad/s; 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003678_a:1008896010368-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003678_a:1008896010368-Figure6-1.png",
        "caption": "Figure 6. Reference trajectory for the lifted leg.",
        "texts": [
            " The encoder\u2019s signal was not sufficient for deriving the value of the angular velocity using a conventional F/V converter, so the value was instead calculated from measurement of the pulse width of the signals. Basic specifications of the experimental biped are shown in Table 1. In the experiments, the biped robot was controlled to walk at a constant speed with a fixed trunk height. The reference for the trunk height was given such that the robot does not go through the singular configurations. The reference trajectory for the lifted foot is shown in Fig. 6. The step length is 140 [mm], and the length of the foot is 50 [mm]. Figure 7 shows the response of the position of the trunk. Figure 8 shows the constrained forces during the double support phase, in which the forces are estimated using Eqs. (25) and P1: VTL/TKL P2: EHE/TKL P3: VTL/TKL QC: PMR/TKJ T1: PMR Autonomous Robots KL465-06-Mitobe May 16, 1997 17:19 294 Mitobe et al. H(q) =  m1 ( a2 1 + l2 1 )+ 2m2l2 1 + m3l2 1 + I1 (m1l2 + m2(a2 + l2)+ m3l2)l1c1\u22122 (m1l2 + m2a2)l1c1\u22123 m1a1l1c1\u22124 (m1l2 + m2(a2 + l2)+ m3l2)l1c1\u22122 m1l2 + m2 ( a2 2 + l2 2 )+ m3l2 2 + I2 (m1l2 + m2a2l2)c2\u22123 m1a1l2c2\u22124 (m1l2 + m2a2)l1c1\u22123 m1a1l1c1\u22124 (m1l2 + m2a2)l2c2\u22123 m1a1l2c2\u22124 m1l2 2 + m2a2 2 + I2 m1a1l2c3\u22124 m1a1l2c3\u22124 m1a2 1 + I1  "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003996_37.648630-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003996_37.648630-Figure2-1.png",
        "caption": "Fig. 2. Regions of the switchingplane for a nonlinear system (Andronov and Maier",
        "texts": [
            " By the micl- 1940s Anclronov was irivcsrignting thc liighcr-order nonlincar systems associated M-ith control engineering, beginning with third-order systcins ivhich ;ire linear except foroiic nonliiiearity caused by arelay or tp C'oulonib friction. His \"point transforii!alion\" inethod. firsi puhlishctl in 1934 [ 121, is of pnrficiiliir interest. Andlont iiid collc,agiies made ii rigorom study of stability of such ,)/stems by scarchiiig [or fixctl points of tr.ans~oriiialioils or thc switching plmc (Fig. 2). The lechnique its tlcvcloped by hitlronov arid colleagues in the 1940s is a direct descendent of hndronov's own late 1920s work on linii l cycles, a i d was gradually extended to hifiher-dimensio~~ state spaczs. and gathered further information on the emerging discipline. The conference was held at a time of intense scientific andpoliticalcriticismof the work being carried out at the new Institute of Automation and Remote Control, the precise circumstances of which are currently the subject of further research by the present author"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002373_j.biosystems.2020.104187-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002373_j.biosystems.2020.104187-Figure4-1.png",
        "caption": "Figure 4. Membership functions for the fuzzy output variables of ,- and \u2205\u2206",
        "texts": [
            " As indicated in Equation (3), M2N/O is the median value of the angles at which the focus swarm robot detects the neighboring robots. For , , denotes the number of the focus robot whereas [ ] = [ \u2026 \u2026 ] represent the numbers of the detected neighboring robots. M2N/O = \"P&Q K , L (3) Fuzzy antecedent variables of , and M2N/O are evaluated respectively with the universal sets of [0 5] and [\u2212 ] and fuzzy sub-sets of R, S, T(L: Low , S: medium, T: High) as shown in Figure 3. Figure 3. Membership functions for fhe fuzzy antecedent variables of , and \u2205M2N/O As shown in Figure 4, the output variables are determined as ,- and \u2205\u2206 subject to the fuzzy antecedent variables of , ve M2N/O as shown in Stage-1. ,- and \u2205\u2206 output variables are determined subject to the R, S, T membership functions in fuzzy universal sets of [0 1] and [\u2212 ] respectively. ,- and \u2205\u2206 indicate the speed and angle of rotation for the robot on GRF. The ,- and \u2205\u2206 fuzzy output variables are determined in Stage-1 for the , and M2N/O fuzzy antecedent variables via the fuzzy logic rules presented in Table 2. In the FLC in Stage-1, FIS(fuzzy interface systems) type Mamdani, rule connection \u201cQ P\u201d method, \u201c & \u201d and \u201c<\" ) $&P\u201d for defuzzification are used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000506_s11771-016-3101-5-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000506_s11771-016-3101-5-Figure1-1.png",
        "caption": "Fig. 1 Circular profile blade and generating revolution surfaces for gear head-cutter: (a) Illustration of circular profile blad; (b) Generating tool surface for convex side; (c) Generating tool surface for concave side",
        "texts": [
            " In no-generating (formate) processes, the generating cradle axis is fixed, no generating roll is employed, so the profile shape of tooth on a workpiece is produced directly from the profile shape on the tool. 2.1.1 Part (a) of generating surface for gear circular profile head-cutter The geometry of the circular profile blade for the gear head-cutter is represented in this section. Figures 1(a), (b) and (c) show illustration of the circular profile blade for the gear head-cutter, and the generating tool surfaces for the convex and concave sides, respectively. As shown in Fig. 1, the coordinate system S0{X0, Y0, Z0} is rigidly connected to the gear head-cutter; Z0 is the axis of the head-cutter; \u03b82 is the rotation angle; each side of the blade generates two sub-surfaces denoted as parts (a) and (b) of the generating surfaces; the segment of the circular arc (part (a)) with the curvature radius R2 (including the curvature radius of the concave side R2c and the convex side R2d) generates the working part of the gear tooth surface; the circular arc of radius \u03c12 generates the fillet of the gear tooth surface (part (b)); \u03b12 is the pressure angle (including the pressure angle of the concave side \u03b12c and the convex side \u03b12d) at the reference point M; s02 is the distance from the top of the blade to the reference point; Pw2 is the point width; r02 is the cutter mean radius; r2 is the cutter point radius (including the pressure angle of the concave side r2c and the convex side r2d); s2 and \u03b82 are the surface coordinates of the part (a); \u03bb2 and \u03b82 are the surface coordinates of the part (b)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000418_978-3-662-46466-3_16-Figure16.3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000418_978-3-662-46466-3_16-Figure16.3-1.png",
        "caption": "Fig. 16.3 Two-link robot",
        "texts": [
            " Evolution over, output optimization solution Based on the optimal individual which was operated by artificial selection, using forward kinematics algorithm, calculate the end location-posture of the robot. If the position is less than E-position and gesture is less than E-gesture Get an inverse kinematics solution of the optimal results. Else Go to Step 2. The effectiveness and performance of the presented algorithms are verified by the experiment (two-link robot and TA1400 robot) and comparative analysis. Simulation experiments were performed on a two-link robot shown in Fig. 16.3. The inverse kinematics problem for the two-link robot can be stated as min a1c1c2 a1s1s2 \u00fe a2c1\u00f0 \u00de2 x2 \u00fe a1s1c2 \u00fe a1c1s2 \u00fe a2s1\u00f0 \u00de2 y2 ! ; a1 \u00bc 40; a2 \u00bc 30 \u00f016:12\u00de where ci \u00bc cos hi\u00f0 \u00de, si \u00bc sin hi\u00f0 \u00de X, Y denote the numerical values of the Cartesian coordinates at the desired position of the robot. The following control parameters were used for the SGA: Population size = 40, Crossover probability = 0.7, and Mutation probability = 0.3. The control parameters were used for the MPGA is the same for SGA, but mutation probability is determined by Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002037_s00170-020-06366-8-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002037_s00170-020-06366-8-Figure14-1.png",
        "caption": "Fig. 14 Ruptured test specimen",
        "texts": [
            " It should only be used in processes that can add more value beyond that provided by conventional methods [9]. Therefore, there are two main advantages of using the machined substrate hybrid AM strategy for injection mould applications. Firstly, complex design features such as CCC and air vent passages can be fabricated additively. Secondly, downstream post-processing times can be minimised. The result of the tensile test is shown graphically in Fig. 13, and the photo of the ruptured test specimen is shown in Fig. 14. Two possible outcomes were anticipated from the tensile test experiment; rupture occurs either at the bonded interface or the side the material has the lower strength. Rupture at the interface will mean weak fusion bonding. In order to have a better understanding of the result data, tensile test data of stand-alone AM-built heat-treated AlSi10Mg [18] and wrought 5083-H116 [19] are used and listed in Table 4 for comparison. According to the stress/strain curve (Fig. 13), the tensile strength, 0.2% proof strength and the elongation at fracture recorded were 319 MPa, 200 MPa and 7.4% respectively. Moreover, the curve shows that there was no necking or further elongation of the specimen after the maximum tensile stress was reached. Fracture also happened instantaneously straight after, which corresponds to the vertical drop of the stress/strain curve. In Fig. 14, it shows that the position of the rupture appeared slightly beyond the fusion-bonded interface and towards the base substrate side. Further examination of the fracture surface on the broken piece from the substrate side by SEM revealed two regions of materials, \u201cA\u201d and \u201cB\u201d (Fig. 15). From the EDS analysis, the chemical composition of material \u201cA\u201d and \u201cB\u201d corresponds to 5083-H116 alloy (5.4% Mg) and AlSi10Mg (11.2% Si) respectively. It is therefore confident to say that the rupture happened at the substrate side"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001142_cns.2019.8802717-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001142_cns.2019.8802717-Figure1-1.png",
        "caption": "Figure 1. A drone exploiting the jamming signal for range estimation.",
        "texts": [
            " While we acknowledge that\u2014in standard scenario\u2014the distance estimation might be improved using more precise statistical description of the channel, we highlight that the scenario considered in this paper is not standard: we expect the jamming signal to be extremely powerful, and the line-of-sight component being strong respect to the other components. The above assumptions are motivated by the defence strategy: the jamming should be powerful in order to be disruptive at far distances to prevent the drone to get close to the protected area, while line-of-sight will be guaranteed by the jammer itself that is interested to maximise the radiated power (against the drone) by adopting a tracking mechanism. Figure 1 shows a toy example where the drone firstly estimates a received power R0, and subsequently a received power R1 = 2\u00d7R0 (+3dB). Recalling Eq. (1), the drone can estimate to have approached the jammer by a factor d1/d0 = 1/ \u221a 2. Assuming the firmware of the drone has been compromised by a malicious entity (or, simply, that it could be programmed), we can implement a combined navigation system that behaves as usual when the drone receives the radio commands from the remote controller, while it switches to the JAM-ME mode when it detects a jamming attack"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000447_jmes_jour_1960_002_027_02-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000447_jmes_jour_1960_002_027_02-Figure3-1.png",
        "caption": "Fig. 3. Optimum configuration for tube ironing",
        "texts": [
            " The introduction of F3 to the left of b raises the to the left Of ab, showing that p, F1, Tab, and Nab are all J O U R N A L M E C H A N I C A L E N G I N E E R I N G S C I E N C E the shear forces Tab and Tbc increase, then all other forces increase, and hence upper-bound values will be obtained if Tab and Tbc are drawn equal to K(ab) and K(bc) respectively. Different values for the upper bound will be obtained from different positions of the points a, b, and c, and the best upper bound will be that which gives the lowest values of the forces P, F,, and F2. For any given position of b, it can be shown that the best positions of a and c are at the beginning and end of the die contact length, as shown in Fig. 3. There is no direct method Vol2 No 3 1960 at UNIV NEBRASKA LIBRARIES on June 5, 2016jms.sagepub.comDownloaded from however, of obtaining the best position of point b. In Fig. 4a the force polygons of Fig. 2c are combined to give the complete force field for the upper-bound condition. This field may be constructed from Fig. 3 as shown by the dotted lines in Fig. 4a; alblcldl is a parallelogram. Force fields for several other positions of b (Fig. 3) and bl (Fig. 4a) may be superimposed on Fig. 4a and the best or minimum value of the drawing force P obtained by plotting the locus of point el and drawing the appropriate tangent (Fig. 4b). The mean drawing stress ratio p/2K is quickly obtained from this diagram since p/2K = P/2K. h. In the absence of friction the force field is shown in Fig. 4c. For this special case, an analytical solution to the problem may be found. Any position of bl may be denoted by angles + and 0 (Fig. 4c). If I) = 0, then the upper bound is _ - P sin a cos a 2~ - sin (e+ sin (e+,) For general values of I) and 6 f=F{ 1 sin e sin (e--a)+sin and the best upper bound is _ - P I 2K smcr y - -{y+i-2 cos a} wherey = z/ (h/H)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000567_j.jsv.2016.04.020-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000567_j.jsv.2016.04.020-Figure6-1.png",
        "caption": "Fig. 6. Rotational subsystem and its five sub-subsystems.",
        "texts": [
            " The ballscrew is considered as a distributed-parameter system, taking into account its rotational and axial flexibility. The remaining elements, including bearings, coupling, motor rotor, and the sliding component, are all assumed lumped-parameter. Newton-Euler equation and the conventional two-coordinate receptance coupling equation are used in the ballscrew's rotational subsystem and axial subsystem modeling; Multi-rigid-body dynamic modeling, Laplace transformation and complex matrix inversion are used for the sliding component; Eq. (21) are used to connect the three subsystems. The top half of Fig. 6 gives the schematic model of the rotational subsystem, where, d0 and L are the ballscrew's effective diameter and length, and the effective diameter is measured at the root of the thread [4]; l is the nut location at the ballscrew, and the nut location changes with the sliding carriage position YTable, and l\u00bcYTable\u00fe l0; c\u03b8b1 and c\u03b8b2 are the rotational damping provided by the front and rear bearings; kc and cc are the rotational stiffness and damping of the coupling; Jm and cm are the inertia of the motor rotor and the rotational damping provided by the rotor bearings. To circumvent the difficulties coming from complex boundary conditions, the rotational subsystem is divided further into five sub-subsystems, as illustrated in the bottom half of Fig. 6, where, A, B are the front and rear section of ballscrew divided by the nut; C, D are the front and rear bearing block, connected to ballscrew's ends via rotational damping; E is the motor rotor connected to the ballscrew via the coupling. Before connecting, A and B are circular shafts with free ends, and their FRFs can be acquired as [13]: HA\u00f0l\u00de \u00bc ha0a0\u00f0l\u00de ha0al\u00f0l\u00de hala0\u00f0l\u00de halal\u00f0l\u00de \" # \u00bc cot\u00f0\u03bbl\u00de=GIp\u03bb csc\u00f0\u03bbl\u00de=GIp\u03bb csc\u00f0\u03bbl\u00de=GIp\u03bb cot\u00f0\u03bbl\u00de=GIp\u03bb \" # (22) HB\u00f0l\u00de \u00bc hblbl\u00f0l\u00de hblbL\u00f0l\u00de hbLbl\u00f0l\u00de hbLbL\u00f0l\u00de \" # \u00bc cot \u00bd\u03bb\u00f0L l\u00de =GIp\u03bb csc\u00bd\u03bb\u00f0L l\u00de =GIp\u03bb csc\u00bd\u03bb\u00f0L l\u00de =GIp\u03bb cot \u00bd\u03bb\u00f0L l\u00de =GIp\u03bb \" # (23) where, G and \u03c1 are the shear modulus, and the material density; Ip \u00bc \u03c0d04=32; \u03bb\u00bc\u03c9 ffiffiffiffiffiffiffiffi \u03c1=G p : FRFs of C, D, and E are: HC \u00bcHD \u00bc 0; (24) HE \u00bc hm \u00bc 1 Jm\u03c92\u00fe i\u03c9cm (25) The zero FRFs of HC and HD imply that the bearing blocks are assumed to be rigid connected to the machine base and the machine base receptance was ignored here"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.20-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.20-1.png",
        "caption": "Fig. 10.20 Schematic diagram of circular equipment",
        "texts": [
            " The broken NC naturally falls into the water flushing device and flush to the washer with a large amount of water. The main drawbacks of centrifugal acid removal equipment include contamination of driven waste acid with NC particles, cloud in waste acid, NC decomposition accident during the acid removal process, acid leaking in the outlet, significant acid consumption, and so on. To reduce acid consumption, diluted acid and water can be used for washing and replacement. (2) Circular equipment The main structure of the circular equipment is shown in Fig. 10.20. The circular equipment has an annular shell of 7800 mm in diameter, and the inner annular groove is driven by an electric motor. The revolving body consists of 20 blocks of movable panels. The fan-shaped plate has some holes with 3 mm in diameter, which facilitates the filtration of acid water. The base of circular equipment is equipped with a variety of bottom tanks for the recovery of different acids, in which some baffles are used to separate the various acids, and the collected acid flow into the recycle bottom tank through bottom collecting pipe"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001978_0954407020964625-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001978_0954407020964625-Figure13-1.png",
        "caption": "Figure 13. (a) Bearing fault implantation rig and (b) four fault bearings.",
        "texts": [
            " Figure 12(d) is the learning result obtained by L1/2-SF, it is observed that the feature trends of the three samples are more consistent than that of L1-SF, especially at the peak position of the 60th feature point. Even the feature amplitude of 1500 r/min sample is perfectly constrained and highly consistent with the other two samples. Thus L1/2-SF can overcome the amplitude and frequency changes caused by variable rotational speed, and the learned feature vectors are more conducive to fault identification and classification. Case 2: Fault diagnosis of a motor bearing The bearing fault implantation rig is displayed in Figure 13(a). The rig contains a motor, a driving belt, a shaft coupling and a bearing seat. The type of the sensor is PCB353B33 made by PCB Piezotronics, Inc, and the sampling frequency is 25.6 kHz. The type of the bearing is NU205EM cylindrical roller bearing, and the designation is wire cutting. There are five bearing health conditions: normal condition (NC), inner ring fault (IF), outer ring fault (OF), roller fault (RF), concurrent fault in outer ring and roller (ORF). Figure 12(b) displays the four fault bearings, and the variable rotating speed information of all bearing fault types is shown in Figure 14"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002764_0142331221994393-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002764_0142331221994393-Figure5-1.png",
        "caption": "Figure 5. Emergency obstacle avoidance strategy.",
        "texts": [
            " The existence of dynamic obstacles affects the motion of the car-like robot. The velocity of the robot gradually decreases along with a gradual increase in the dynamic potential field function. The coefficients of the three potential field functions determine their weight proportions in the final APF. For example, to improve the motion timeliness of the car-like robot, the coefficient of the attractive potential field should be increased appropriately. Controlling the steering angle and acceleration of the carlike robot in real time is necessary. As shown in Figure 5(a), the detection and buffer areas are set in a dangerous scenario where dynamic obstacles suddenly appear. The detection area refers to the area where the car-like robot can pass with a maximum steering angle, whereas the buffer area refers to the expansion range of the detection area that ensures that the car-like robot will not collide with the obstacles. Both areas constitute a safe area for the movement of the car-like robot. In this study, two situations are considered, namely, the entrance and exit of dynamic obstacles in the safe area. Figure 5(b) shows that when a dynamic obstacle enters the safe area, the obstacle and the car-like robot constantly approach each other, the dynamic and repulsive potential field functions increase, and the velocity of the car-like robot decreases. If the planned path that bypasses the obstacle is used to prevent the car-like robot from hitting the obstacle, then this path is not smooth, and the movement distance subsequently increases. This phenomenon does not satisfy the constraints of the TEB approach. As shown in Figure 5(c), when a dynamic obstacle leaves the safety area, the functions of the dynamic and repulsive potential fields decrease, and the velocity of the car-like robot gradually increases. At this time, a sharp turn or stop is not necessary to avoid the obstacle. The TEB approach can generate a reasonable and smooth local path. The function defined in Equation (20) is introduced to control the acceleration and steering angle. When the dynamic obstacle is in the safe area, the relative movement trend between the dynamic obstacle and the car-like robot is evaluated by the dynamic potential field"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000524_s0005117916010069-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000524_s0005117916010069-Figure3-1.png",
        "caption": "Fig. 3. The problem geometry. On the left: control includes pitch \u03b2(t) and yaw \u03b3(t) angles. On the right: observed azimuth \u03d5A and elevation \u03d5E angles.",
        "texts": [
            " In this section, by an observer P we mean UAV that can perform 3D space maneuvers for trajectory control of observations under given constraints, controlling pitch and yaw angles. The AUTOMATION AND REMOTE CONTROL Vol. 77 No. 1 2016 target E, similar to Section 2, moves on a plane uniformly along a straight line. We can observe (with noise) its azimuth and elevation angles. To find out how the accuracy of TME estimates is influenced by UAV\u2019s spatial maneuvering, in Section 3 we consider the TME estimation problem only. Let us consider the Cartesian coordinate system OXY Z (Fig. 3) whose OX, OY , and OZ axes point to east, north, and up respectively. We assume the Earth surface to be locally flat and define it with equation Z = 0. Suppose that along this plane, the target moves uniformly along a straight line with speed v. Consider the relative coordinate system oxyz with axes parallel to the OXY Z system whose origin is at UAV\u2019s center of mass. Initial coordinates of the target (at the time moment t = 0) in this system equal (x0, y0, z0), where z0 is the initial altitude of UAV\u2019s flight taken with an opposite sign",
            " Motion equations for UAV\u2013target system have the following form: \u23a7 \u23aa\u23a8 \u23aa\u23a9 x\u0307(t) = vx \u2212V cos \u03b2(t) cos \u03b3(t) y\u0307(t) = vy \u2212V cos \u03b2(t) sin \u03b3(t) z\u0307(t) = \u2212V sin\u03b2(t), (10) where \u03b2 = \u03b2(t) and \u03b3 = \u03b3(t) are UAV controls with respect to pitch and yaw angles respectively, and V = const is a given speed of UAV\u2019s flight. We will assume that at the initial time moment t0 = 0 we have a prior estimate \u03020 for the TMEs, and we know the 4\u00d7 4 covariance matrix of the estimation errors for P0. Evolution of matrix Pt is defined in the Appendix. The UAV can observe the column vector \u03bet = ( \u03beA(t), \u03beE(t) ) (here \u03beA(t) and \u03beE(t) are noisy azimuth and elevation angles respectively) with passive bearing means in either radio [17] or infrared range (Fig. 3). Such an observation scheme has been presented, for instance, in [16]. Suppose that at every time moment coordinates of UAV itself, the horizon line, and the direction to the north are known exactly, and a detecting device does not change orientation in the UAV maneuvers, i.e., there are no systematic errors in observations. Under these assumptions, the observation equation has the form \u03bet = ( \u03d5A(t) \u03d5E(t) ) + wt, where wt = ( \u03c3Aw1(t) \u03c3Ew2(t) ) , w1(t) and w2(t) are independent standard Wiener processes, \u03c3A and \u03c3E are given constants"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.34-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.34-1.png",
        "caption": "Fig. 10.34 The flow of the NC slurry in the chopper",
        "texts": [
            " After the introduction of fine-breaking technology to the NC manufacturing, the stability issue of NC has been solved, which not only achieves a new milestone of NC production history but also realizes the wide application of NC [2, 7]. Fine-breaking is the cutting and grinding of the long fiber of NC after boiling by the physical and mechanical action of the chopper to increase the surface area of NC, which facilitates the diffusion of residual acid and other unstable materials in the capillary, thus accelerating NC stability treatment. The flow of the NC slurry between the rolling cutter and the bed knife is shown in Fig. 10.34 in the process of NC fine-breaking. Affected by centrifugal force, friction resistance, and other forces, NC slurry is strongly impacted and mixed. The flow of the NC slurry is divided into two parts by the rolling cutter. One part flows upward along the front wall of the rolling cutter, whereas another part flows forward through the bottom of rolling cutter. The NC slurry flowing into the blade flute forms eddy motion within the groove under the effect of centrifugal force and friction, which is illustrated by the formation of a groove on the front edge of the knife after a long time use"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002446_j.amc.2020.125609-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002446_j.amc.2020.125609-Figure3-1.png",
        "caption": "Fig. 3. (a) Meshing and (b) post-processing CFD analysis of a modular block.",
        "texts": [
            " In order to completely define the flow a turbulence k- model [8] is used for being the most suitable one for the turbulence levels and Reynolds number values that are reached inside this sort of structures. The resolution numerical scheme used in resolving the problem is the \u201cpressure based\u201d type, using a \u201csimple\u201d type scheme for the coupling between pressure equations and velocity ones. Based upon this model configuration and input flow definition, the analysis is performed. Simulation runs for \u201ctransient regime\u201d and it concludes when \u201cpermanent regime\u201d is established in the system. Once calculation process is completed one proceeds to results post-processing Fig. 3b, that allows to analyse the flow characteristics developed within modular structures as well as to determine depth profiles along the models studied. From which will be possible to obtain Manning coefficient (n) and head minor losses dimensionless (k) values, these parameters characterize energy dissipation produced by the mentioned modular elements existence. Once CFD model has been built is necessary to define the case studies will proceed to address. Since the aim pursued is to determine energy dissipation of these modular blocks in both directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003659_mssp.1998.0190-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003659_mssp.1998.0190-Figure7-1.png",
        "caption": "Figure 7. The four worst selections of excitation conditions: (a) all the excitations meet at one point; (b) all the extension lines of excitations meet at one point; (c) all the lines of excitation are parallel; (d) extension of excitation line passes through the centre of gravity.",
        "texts": [
            " Therefore, in order to determine the moments of inertia accurately, the excitation direction/points should be selected so that the magnitude to volume ratio of the angular acceleration vectors is as small as possible. Needless to say, this condition can be accomplished when the angular acceleration vectors are mutually orthogonal to one another. 5.3. THE WORST EXCITATION CONDITIONS Upon investigating equation (28) carefully, the worst excitation conditions can be derived as follows: Case 1: All the excitation points meet at one point, as shown in Fig. 7(a), or all the extension lines of excitations meet at one point as shown in Fig. 7(b). Proof. When the moment of inertia matrix about the centre of gravity is de\"ned [J],C J xx !J zy !J xz !J xy J yy !J yz !J xz !J yz J zz D (34) the angular acceleration vector Mq j N can be written Mq j N\"[J]~1MMg j N\"[J]~1[\u00b9e j ]MF j N . (35) Since [\u00b9e 1 ]\"[\u00b9e 2 ]\". . .\"[\u00b9e m ] in these cases, matrix [Q] can be written as [Q]\"[Mq 1 N, Mq 2 N, . . . , Mq m N]\"[J]~1 [\u00b9e 1 ] [MF 1 N, MF 2 N, . . . , MF m N]. (36) Recalling that [\u00b9e 1 ] is the 3]3 skew-symmetric matrix and the determinant of [\u00b9e 1 ] is zero, the determinant of [Q] [Q]T becomes identically zero. Det ([Q] [Q]T) \"(Det ([J]~1))2 Det([\u00b9e 1 ] [\u00b9e 1 ]T) Det ([MF 1 N, . . . , MF m N] [MF 1 N,2 , MF m N]T)\"0. (37) Therefore, in these excitation conditions, the rank of [R q ] is always de\"cient. Case 2: All the lines of action, as shown in Fig. 7(c), are parallel. Proof. Since all the excitation directions are the same, the applied force vectors can be written as F j \"c j (F x i#F y j#F z k)\"c j F 1 , j\"2, 3, . . . , m (38) where c j is a constant scalar. When moment vectors MM x N, MM y N, MM z N are de\"ned MM x N\"[Mx 1 Mx 2 . . . Mx m ]T MM y N\"[My 1 My 2 . . . My m ]T MM z N\"[Mz 1 Mz 2 . . . Mz m ]T (39) the matrix [Q] can be written as [Q]\"[J]~1[MM x NT, MM y NT, MM z NT] \"[J]~1[[\u00b9e 1 ]MF 1 N, c 2 [\u00b9e 2 ]MF 1 N, . . . , c m [\u00b9e m ]MF 1 N]. (40) In this case, the moment vectors MM x N, MM y N, and MM z N are linearly dependent, since F x MM x N#F y MM y N#F x MM z N\"0. (41) Hence, the rank of [Q] is always de\"cient. Therefore, the determinant of [R q ]T[R q ] is identically zero. Case 3: As shown in Fig. 7d, the extension line of excitation passes through the centre of gravity. Proof. Since angular accelerations in this excitation condition are identically zero, the angular accelerations due to this excitation cannot contribute to the rank of [R q ]. 5.4. SUGGESTIONS FOR THE SELECTION OF EXCITATION DIRECTION/POINTS The excitation conditions should be selected so that the area and the volume of the angular acceleration vectors are as large as possible. Nonetheless, the selection of excitation conditions for the volume of the acceleration vectors is more important than that for the area"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001671_j.compstruct.2020.112468-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001671_j.compstruct.2020.112468-Figure1-1.png",
        "caption": "Fig. 1. Features of a typical composite T-joint.",
        "texts": [
            " Co\u2010curing integration of many substructures during manufacturing is easy and the assembly time and cost are eliminated. The integration of several parts together have other added advantages of reduced part count, simplified assembly process, reduced manufacturing cost, smooth aerodynamic contour, smooth load transfers etc. The basic element of co\u2010curing construction is the T\u2010joint, which forms the primary load transfer mechanism between skin and stiffener. The individual element of T\u2010joint is referred in Fig. 1. The web provides an interface for attachment of the skin and radius fillet provides continuity of load transfer between web and flange. This kind of T\u2010joint is generally referred to as normal/plain T\u2010joint. Nearly all studies found in literature deal with this configuration for the evaluation of joint behaviour. In a normal T\u2010joint, a stiffener is bonded to the skin. The stiffeners could be of various shapes, such as, T, I, C, L, J etc. However, at the joint, all the configurations are similar to Fig. 1. Composites possess very good in\u2010plane properties but have poor out\u2010of\u2010plane characteristics [1]. When T\u2010joints are subjected to out\u2010 of\u2010plane loading, the joint typically fails at skin flange bond interface [2]. Many researchers have addressed this issue by evaluating the bond interface strength [3\u20137]. Several methods have been suggested in literature to improve out\u2010 of\u2010plane properties of composites manufactured using prepreg/autoclave molding process. A few of them are, use of stronger adhesive [8], tough thermoplastic film [9] and nanoparticles at interface [10]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001749_s12555-019-0922-7-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001749_s12555-019-0922-7-Figure1-1.png",
        "caption": "Fig. 1. Heterogeneous multi-agent formation graph.",
        "texts": [
            " The graph G is allowed with several components, and every components are connected via undirected edges in some topologies. The graph G\u0304 of this MAS is said to be connected if at least one agent in each component of G is connected to the leader by a directed edge [8]. Let Ni be the set of neighbours of agent i. Moreover, ai j = a ji(i = 1, ...,n; j = 1, ...,n) with aii = 0 denotes the nonzero interconnection weights between agent i and agent j. gi > 0 if there is a link between the leader and the follower i. For example, in Fig. 1, a24 means the interconnection weights between agent 2 and agent 4, g4 denotes the link weight value between the leader and the follower 4. Then, the Laplacian of the weighted graph G is Distributed Coordination of Heterogeneous Multi-agent Systems with Dynamic Quantization and L2\u2212L\u221e Control 3 denoted by L. E is an in-degree matrix that E = diag(di) with di = n \u2211 j=1 ai j. In this paper, the coordinated object of the heterogeneous MAS formation mainly focuses on planar motion, namely, the UAV formation is expected to be able to track the UGV from a plane perspective",
            " Assume that the leader is active in the ground and transfers the system dynamics to followers, the dynamics discrete-time system of the leader is described as follows [31]: l0(k+1) = Sl0(k)+ r(k), y0(k) = Rl0(k), (2) where l0(k) \u2208 Rp is the system state, and y0(k) \u2208 Rm is the output. r(k) is the external bounded signal. S and R are the matrices with appropriate dimensions, and S is Hurwitz stable. For the heterogeneous MAS, the formation of the followers is specified by a vector h(k) = [hT 1 (k), hT 2 (k), ..., hT n (k)] T , where hT i (k) \u2208 Rm and hi(k) is bounded. For instance, in Fig. 1, the UAV 1 tracks the UGV leader and maintains a formation with vector h1(k), and heterogeneous MAS in this paper will be extended to n agents with hn(k). The control objective of this paper is to design the distributed control scheme ui in (1) for each follower to ensure that lim k\u2192\u221e \u2016(yi(k)\u2212hi(k))\u2212 y0(k)\u2016 = 0 in the given formation. The output adjustment error is defined as follows: \u03b7i(k) =y0(k)\u2212 (yi(k)\u2212hi(k)) =Rl0(k)\u2212 (Cxi(k)\u2212hi(k)) . (3) Remark 1: For the given formation, the formation vector hi(k) 6= 0, which will illustrate in the following assumptions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001966_j.mechmachtheory.2020.104136-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001966_j.mechmachtheory.2020.104136-Figure8-1.png",
        "caption": "Fig. 8. (a) Gear set model. (b) Equivalent model of gear set.",
        "texts": [
            " In recent literature on driveline modeling, the gear sets are assumed as rigid elements and the gear box is considered as a lump element, whose inertia is calculated based on the gear ratio. In fact, when the gear set transmits torque by teeth meshing, the number of meshing teeth varies with the rotation speed of the gears, which further results in varying contact stiffness of the gears. The gear set meshing and the contact stiffness are shown in Fig. 7 . With the assumption that the bearings, which support the gear shaft are rigid, and that the axial force of the helical gear set is not considered, the gear set model can be simplified to have only torsional motion issues. Fig. 8 presents the diagram of the gear set motion and the equivalent torsional motion model. The dynamic model of the gear set can be obtained based on the Newton\u2019s second law, and can be expressed by Eq. (8) . { J i \u0308\u03b8i + F d r i \u2212 T i = 0 J k \u0308\u03b8k \u2212 F d r k + T k = 0 F d = k mik ( t ) d is ( t ) + c mik ( t ) d is ( \u02d9 t ) dis ( t ) = r i \u03b8i \u2212 r k \u03b8k \u2212 e ik dis ( \u02d9 t ) = r i \u02d9 \u03b8i \u2212 r k \u02d9 \u03b8k \u2212 \u02d9 eik (8) Where, r i, r k is the base circle radius of the gear i and k, respectively. k mik ( t ) is the contact stiffness of the gear set in the torsional direction, k mik ( t ) = cos \u03b2k m ( t ), k m ( t ) is the mesh stiffness of the helical gear,which is referred to as the timevarying mesh stiffness, \u03b2 is the helical angle of the gear set"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000178_s10015-019-00564-8-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000178_s10015-019-00564-8-Figure1-1.png",
        "caption": "Fig. 1 Underwater vehicle-manipulator system",
        "texts": [
            " The controller proposed in this paper is the redesigned version of the controllers developed in\u00a0[22, 23]. The controller designed in\u00a0[22] can be applied only to a torque-controlled servo system, whereas that in\u00a0[23] is applicable only to a velocity-controlled one. On the other hand, the proposed controller consists of three sub-controllers, and its control input type can be changed by means of the combination of the sub-controllers. An underwater vehicle-manipulator system (UVMS) considered in this paper is shown in Fig.\u00a01. It consists of the robot body (the vehicle) and the NL link manipulator with revolute joints. The vehicle is propelled by marine thrusters, whereas the manipulator is directly driven by electric motors. When the UVMS can move in a (3D) space, we have NL = 3 . On the other hand, when its motion is restricted in a (2D) plane, we have NL = 2 . Without loss of generality, we will show only results of the 3D motion analysis (i.e., NL = 3 ). It should be noted that when we choose the third element of a vector for linear motion (e.g., a position vector) and the first and second elements of a vector for rotational motion (e.g., an orientation vector) to be zeroes, we can perform a 2D motion analysis in a way similar to a 3D motion analysis. The meanings of the symbols shown in Fig.\u00a01 are explained in Table\u00a01. Mathematical models of the UVMS shown in Fig.\u00a01 can be represented as a series connection among a kinematic model, a dynamic one, and an actuator\u2019s, as illustrated in Fig.\u00a02. The equations of the blocks will be derived below. First, the kinematic model used in this paper is expressed as The model (1) is represented with respect to the vehicle coordinate system V , and hence, it is the same as the model of a terrestrial manipulator (pp. 149\u2013151 in\u00a0[24]). Although (1)p\u0307V M (t) = JM(q)q\u0307(t). 1 3 the model (1) will be used in the controller design, the stability property of the position with respect to the inertial coordinate system I [i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001834_j.addma.2020.101547-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001834_j.addma.2020.101547-Figure6-1.png",
        "caption": "Fig. 6. Schematic diagram of the fixture.",
        "texts": [
            "0mm) in the DED process is shown in Fig. 4: hatch spacing was w=2.0mm and there were 11 fusion lines in each layer. Figs. 4a and 4b show the deposition tracks of the first and second layers, respectively. As shown in Fig. 4c, the deposition orientations of the first and second layers were perpendicular to each other. The substrate used in this study is a thick substrate with a thickness of 16mm, which is made of 304-stainless steel. As shown in Fig. 5, the bottom of the substrate was polished. As shown in Fig. 6, a special fixture is designed in the experiment. The substrate was coaxially clamped on the center hole of the support frame by two bolts. The fixture consists of bolts, support frame and springs. The gap between the substrate and the fixture ensures that the substrate can move freely in the x- and y-directions. The substrate is only constrained by the spring force in the z-direction. Fig. 7 shows the entire experimental setup for in-situ monitoring using the CGS system in the experiment. In this study, we performed in-situ monitoring for measuring substrate deformation in the DED process using a CGS system to investigate the substrate deformation characteristics at different stages"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002837_taes.2021.3069035-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002837_taes.2021.3069035-Figure2-1.png",
        "caption": "Fig. 2. Schematic diagram of line deviation measurement.",
        "texts": [
            " First of all, the effective range of spatially coded laser beam named guidance field is shown in Fig. 1. The guidance field is created and adjusted by the LBPU, so that it diverges quickly at the beginning to chase the missile over a wider range and fixes to a constant after that to keep a proper signal strength. The beam centerline directly points to the target and guides the missile to fly along it until hitting the target. The line deviation of missile to the beam centerline is inevitable. As shown in Fig. 2, the line deviation is measured by modulation code in spatially coded laser beam, and modulation coding is completed by the modulation disk. When the modulation disk works, it rotates clockwise, and the inner and outer code channels staggered by 90\u25e6 are used to measure the longitudinal and lateral line deviations, respectively. Thus, the main task of guidance and control system is to keep the missile within the guidance field and eliminate the line deviation. Note that the laser beam riding missile is usually axisymmetric and controlled by skid-to-turn scheme, and thus the longitudinal and lateral models have the same structure",
            " The expressions d2 = D1, d3 = qS mg c \u03b1 yD2 and d4 = D3 where D1, D2 and D3 are represented as D1 = \u2206L m + \u2206D1, D2 = \u2206L mV + \u2206D2 and D3 = \u2206M Jz + \u2206D3. Note that the transformation between LBPU constraints and controlled output is designed in (6), and the limitation of practical measurement is considered in (7) and (8). Therefore, an improved IGC model in the presence of mismatched unknown uncertainties is finally established. Remark 4: In (8), x1 is measured by modulation code in spatially coded laser beam as shown in Fig. 2, when the modulation coding is completed by the modulation disk; x2 is obtained by integrating the acceleration measured by accelerometer; x3 and x4 are measured by the accelerometer and gyroscope mounted on the laser beam riding missile, respectively. IV. IGC LAW DESIGN The EDO to deal with mismatched unknown disturbances is introduced to estimate the uncertainties for the improved IGC model, then the SMC is applied to design the IGC law. A. Introduction of the EDO Assumption 1: [29]\u2013[32] In the IGC system (8), the mismatched unknown uncertainties di(i = 2, 3, 4) and their derivatives are bounded by known positive constants, satisfying that |d(j) i | \u2264 \u03c8i, i = 2, 3, 4, j = 0, 1, 2",
            " degrees in navigation guidance and control both from Northwestern Polytechnical University, Xi\u2019an, China, in 2013 and 2017, respectively. He is currently an Associate Researcher of navigation guidance and control with the School of Astronautics, Northwestern Polytechnical University, Xi\u2019an, China. His research interests include robust control, sliding mode control, coupling control with their applications to aerospace systems. FIGURE CAPTION Fig. 1. Schematic diagram of guidance field for laser beam riding guidance. Fig. 2. Schematic diagram of line deviation measurement. Fig. 3. Schematic diagram of the proposed control scheme. Fig. 4. State response curves in the presence of external disturbances under EDO-FTPPF-IGC, FTPPF-IGC and IGC. Fig. 5. Input curves in the presence of external disturbances under EDO-FTPPF-IGC, FTPPF-IGC and IGC. Fig. 6. Disturbances, estimations and estimation errors of EDO. Fig. 7. State response curves with T0 = 3s, 3.5s, 4s and 4.5s under EDO-FTPPF-IGC. Fig. 8. Input curves with T0 = 3s, 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001837_isitia49792.2020.9163758-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001837_isitia49792.2020.9163758-Figure3-1.png",
        "caption": "Fig. 3. Differential mobile robot.",
        "texts": [
            " Single robot deployed in this study has diagram block depicted in Fig. 2. The task can be managed by sensors attached in robots. Two gas sensors as stereo noses are used to navigate the robot that tracks gas plume toward its origin based on the gas concentration \u03b2l, \u03b2r. LiDAR will locate other robots location as polar coordinate d, \u03b8 and navigate robot to maintain distance between robots so that the formation can be performed. Robot deployed in this study is differential steering mobile robot that has two actuators that handle robot navigation as shown in Fig. 3. Fuzzy logic relates between sensors as input and actuators as output while performing defined tasks. Velocity v of robot as output can be derived by angular velocity of each wheel \u03c9l, \u03c9r multiplied by the radius of the wheel R as presented in (1). Then angular velocity can be derived by multiplying the angular velocity of each wheel with the radius of the wheel and the width of the robot body l. = 2 2\u2212 (1) Action of robot in behavior-based formation control can be derived into subtask depicted in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure2-1.png",
        "caption": "Fig. 2. Initial pose of 3-R(RRR)R+R HAM.",
        "texts": [
            " The 3-R(RRR)R parallel manipulator consists of a fixed platform, a moving platform and three R(RRR)R chains. To adapt to the requirement of non-circular polarization antenna phase invariance, the polarization mechanism is connected in series between the 3-R(RRR)R parallel manipulator and the antenna reflector, so that the antenna reflector can rotate around the central axis of the moving platform. The novel 3-R(RRR)R+R HAM has 3R1T 4-DOF. Figure 1 shows the geometrical diagram of 3-R(RRR)R+R HAM. 3-R(RRR)R+R HAM is shown in Fig. 2. The fixed and moving platforms are expressed as P1 P2 P3 and Q1 Q2 Q3, respectively, and the branch chains are expressed as Pi Ji Gi Ki Qi (i = 1, 2, 3). The fixed coordinate system O \u2212 Ox Oy Oz and the moving coordinate system C \u2212 CxCyCz are, https://doi.org/10.1017/S0263574720001290 Downloaded from https://www.cambridge.org/core. University of Toledo, on 03 Jun 2021 at 19:25:40, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. respectively, established at the center of the fixed and moving platform",
            " Altogether, k, \u03c2 and r are the pitch angle, azimuth angle and relative fixed coordinate distance of the dynamic coordinate system of the antenna, respectively. The angle \u03b86 between Tz-axis and Cx -axis represents the polarization angle. In the process of antenna tracking satellite, pitching and azimuth motion are usually required. The pitching and azimuth motion process of the 3-R(RRR)R+R HAM is analyzed. During the pitching motion, the azimuth of the antenna remains unchanged. The initial pose of the 3-R(RRR)R+R HAM is the initial pose for antenna movement, as shown in Fig. 2. The expected pose of the pitch motion is shown in Fig. 3. In the initial pose, a certain vector T V is given in the end coordinate system Tx T Ty and the direction of the vector is along the intersection line of plane OCT and plane Tx T Ty . The direction of vector T V is always along the intersection of plane OCT and plane Tx T Ty , and the rotation angle of the polarization mechanism is \u03b86 = 0\u25e6. Therefore, in the process of pitching, not only the antenna reflector does not move with it but also the polarization mechanism remains stationary",
            " The HAM moves from the initial pose to the transition pose I and transition pose II to the expected pose, and the HAM does not rotate when it is pitching; when the HAM moves from the transition pose I to transition pose II, the polarization mechanism rotates. The 3-R(RRR)R+R HAM can solve the problem of the phase change of the reflector caused by the accompanying motion by introducing a single degree of freedom polarization mechanism and meet the working requirements of non-circular polarization antenna. As shown in Fig. 2, the 3-R(RRR)R+R HAM is in the initial pose and the position vector of revolute joints Pi = (i = 1, 2, 3) in the coordinate system O \u2212 Ox Oy Oz can be expressed as https://doi.org/10.1017/S0263574720001290 Downloaded from https://www.cambridge.org/core. University of Toledo, on 03 Jun 2021 at 19:25:40, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. 6 Dynamic modeling and mobility analysis of the 3-R(RRR)R+R antenna mechanism\u23a7\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23a9 \u25e6P1 = [R, 0, 0]T \u25e6P2 = [ \u2212R/2, \u221a 3R/2, 0 ]T \u25e6P3 = [ \u2212R/2, \u2212\u221a 3R/2, 0 ]T (1) where R denotes the circumcircle radius of the fixed platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure9-1.png",
        "caption": "Fig. 9. Oil Leakage along the Scoop Tube Housing and Baffle for its Prevention",
        "texts": [
            " The foot valve design was altered and tests made with various materials for the seating, the best results being obtained with a high-grade whitemetal in a mild steel housing. The whitemetal is durable enough to withstand a million reversals of the pump, while it is soft enough to flow very slightly when distortion of the housing occurs, and thus maintain a perfect seal. The cause of the loss of speed and the inconsistency as between one coupling and another was eventually traced to variations in the spray effect in the scoop tube chamber, and the resultant creeping of oil along the scoop tube housing at point X in Fig. 9, and thence down through the manifold drain ports into the reservoir. The leakage is prevented simply enough by fitting an anti-leak baffle as shown at point Y, but its occurrence was only discovered by cutting open the outer casing of a test coupling to study scoop tip effects under various running conditions. In the case of the coupling running clockwise, as shown in Fig. 10a, there was little spray and no need for the anti-leak baffle, since the \u201cbow wave\u201d of oil at the tip of the scoop tube is seen to be quickly merged with the annular ring of oil rotating with the casing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002671_j.jsv.2021.115967-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002671_j.jsv.2021.115967-Figure6-1.png",
        "caption": "Fig. 6. Position and direction of the 12 accelerometers (red arrows) and two lasers (green arrows) on the test-specimen. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " The test-specimen is comprised of an aluminium circular beam with a diameter of 6 mm and a thickness of 1 mm with wooden blocks ( 392 \u00d7 202 \u00d7 27 mm) attached at each end of the beam where one block is clamped to the ground. The position of the second wooden block is adjustable to obtain different lengths between the blocks. We use two different setups where we adjust the length between the wooden blocks so the test-specimen has a height of 762 and 798 mm in the setups. Fig. 5 illustrates the test-specimen. We use 12 accelerometers of the type Br \u022d el & Kj \u022a r 4508-B, 100mV/g and two lasers of type Micro-Epsilon optoNCDT 1300 to measure the wooden block at the top, see Fig. 6 . We excite the structure with a Br \u022d el & Kj \u022a r impact hammer of the type 8206, 22.5mV/N, so we have a free decay of the system. The applied accelerometers are piezoelectric accelerometers where the general working principle behind these are as follows: a seismic mass is attached to a piezoelectric material, when a force act on the mass, the inertia of the mass induces stress onto the piezoelectric material and it produces an electric charge proportional to the force acting on the mass, which is transformed to acceleration through Newtons second law [1,2] ",
            " Using standard experimental modal analysis in a Multiple-Input/Multiple-Output (MIMO) formulation, we apply multiple impact loads to the test-specimen at different locations. We organise all the free decays (0.2 s after each impact) in a matrix, and we estimate the modal parameters from this matrix using a stabilization diagram based on the Ibrahim timedomain method [2,20] , see Fig. 7 . Fig. 8 illustrates the five mode shapes from the identification process and Table 1 holds the identified modal parameters. We create a simple finite element model of the test-specimen simplified to 2D - corresponding to the yz-plane from Fig. 6 - with one translational and one rotational DOF using an Euler-Bernoulli beam element fixed at one end and with a lumped mass with mass moment of inertia in the other end, see the system in Fig. 9 (a). The first and fourth mode from the identification process from the experimental analysis correspond to the given plane. The model is updated to resemble the experimental analysis in terms of natural frequencies, see the modal parameter in Table 1 . Fig. 9 (b-c) illustrates the two mode shapes of the system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000418_978-3-662-46466-3_16-Figure16.1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000418_978-3-662-46466-3_16-Figure16.1-1.png",
        "caption": "Fig. 16.1 Kinematic diagram of Panasonic TA1400 robot",
        "texts": [
            " MPGA divides all populations into several populations, then through artificial selection and immigration operation [7] to form a new population by selecting the best individuals from each category. So the convergence speed of genetic algorithms [8] and inversion precision were significantly improved using this mechanism. Finally, after taking 6 DOF robot TA1400 as an example, the general mechanical structure robot inverse kinematics was solved by using the proposed novel inverse kinematics subproblem solution, and realized intelligent robot trajectory planning. Figure 16.1 is the kinematic diagram of Panasonic TA1400 robot when revolute joint is zero. Kinematic model of robot is built according to the special mechanical structures. i 1 i T is a transformation matrices [9] which is resulted from the coordinate system i \u2212 1 and coordinate system i. The homogeneous transformation matrix for a single joint is expressed as: i 1 i T \u00bc chi shi 0 ai 1 shicai 1 chicai 1 sai 1 sai 1di shisai 1 chisai 1 cai 1 cai 1di 0 0 0 1 2 664 3 775 \u00f016:1\u00de where chi \u00bc cos hi\u00f0 \u00de, shi \u00bc sin h\u00f0 \u00dei, hi is the angle between xi 1 and xi axes measured about zi axis, ai is the angle between zi and zi\u00fe1 axes measured about xi axis, ai is the distance from zi to zi+1 axes measured along xi axis and di is the distance from xi 1 to xi axes measured along zi axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001688_s11661-020-05831-z-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001688_s11661-020-05831-z-Figure1-1.png",
        "caption": "Fig. 1\u2014Schematic diagram of the SLM formation process.",
        "texts": [
            "34, Young\u2019s modulus was set to 195 GPa, and other parameters adopted the default Yade settings. The metal particles are melted by laser heating to form a molten pool during the SLM formation process. The molten pool undergoes intense convection and shape change under the combined effects of surface tension, the Marangoni effect, and gasification recoil. The molten pool then gradually solidifies to form a solidified track and forms a metallurgical bond with the metal substrate or the upper formation layer. Figure 1 shows a schematic diagram of the SLM formation process. To ensure the stability and efficiency of the numerical solution, the following three assumptions were made: ignoring the influence of metal density change on volume, ignoring the influence of metal liquid gasification on alloy composition, and considering the flow of liquid metal and protective gas satisfying the laminar flow condition of an incompressible Newtonian fluid. Describing the SLM formation process based on the mesoscopic scale means that the interface between the metal and the protective gas needs to be characterized in real time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002037_s00170-020-06366-8-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002037_s00170-020-06366-8-Figure5-1.png",
        "caption": "Fig. 5 Layout of mould inserts for printing experiments",
        "texts": [
            " This material is also recommended by EOS, a metal PBF system manufacturer, as their standard material for the build plate for AlSi10Mg powder. Table 2 shows the chemical composition of the 5083-H116 aluminium alloy. All four inserts were fabricated together in a single build on an EOS M290 metal PBF machine. Print preparation, such as honeycomb structure and support creation, were done in Materialise Magics, a commonly used 3D print processing software package [16]. Two print experiments were performed, one using standard practice printed supports, and the other using the proposed hybrid-build method. Figure 5 shows the layout of the mould inserts in both print experiments. As recommended by EOS, the optimised PBF process parameters for AlSi10Mg powder are as follow: \u2022 Laser power: 370 W \u2022 Hatch spacing: 0.13 mm \u2022 Scan velocity: 1300 mm/s \u2022 Layer thickness: 0.03 mm \u2022 Scan pattern: Stripes, 7 mm wide, 67\u25e6 rotation angle In order to achieve maximum relative density from the build and avoid part detachment from the substrate block, the building temperature of the build plate was set as 200 \u25e6C. In order to establish a baseline for time comparison, the four redesigned inserts were first fabricated using standard 5 mm tall solid supports setup (also known as volume supports in Materialise Magics)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002671_j.jsv.2021.115967-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002671_j.jsv.2021.115967-Figure2-1.png",
        "caption": "Fig. 2. As the accelerometer tilts, it measures a component of gravitational acceleration: (a) rotation, \u03b8x , around the x -axis, (b) rotation, \u03b8y , around the y -axis, and (c) rotation, \u03b8z , around the z-axis.",
        "texts": [
            " Thus, we have a triaxial accelerometer mounted in a single point, measuring the translational acceleration in the main directions, see Fig. 1 . In the case of no rotation of the triaxial accelerometer, we express the tilt error as the static tilt error. \u03b5 (t) = g , for \u03b8(t) = 0 (3) where g = [ 0 0 \u2212g ]T is a vector accounting for the gravitational acceleration (sometimes called earth\u2019s gravitational field vector). Next, we will introduce rotation around the x -, y -, and z-axes - denoted as \u03b8x , \u03b8y , and \u03b8z , see Fig. 2 , and we organise these rotations in a vector. \u03b8(t) = [ \u03b8x (t) \u03b8y (t) \u03b8z (t) ]T (4) We express the error as the accelerometer tilts by multiplying with the rotation matrix. \u03b5 (t) = R ( \u03b8(t) ) g (5) where R ( \u03b8(t) ) is composed by the three basic rotation matrices where the order of these rotation matrices depends on the order in which the rotations are applied [6] . By assuming small angles - cos (\u03b8 ) \u2248 1 , sin (\u03b8 ) \u2248 \u03b8 - the order of rotation matrices becomes insignificant and we can choose any order"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001690_s12046-019-1263-1-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001690_s12046-019-1263-1-Figure3-1.png",
        "caption": "Figure 3. Schematic representation of the worm gear test rig.",
        "texts": [
            " The number of nodes in the input layer is equal to the number of features for this network. It is determined as 12 in this study. Statistical parameters of time and frequency domains are used as an input for BPNN. An ANN is composed of two hidden layers and the number of nodes in each hidden layer is determined as 12. The number of output nodes is determined as one. 3.1 Worm gear test rig A schematic representation of the test rig, which consists of an electrical motor, a worm gearbox and an electromagnetic brake, is shown in figure 3. Moreover, the vibration monitoring system consists of the accelerometers, signal conditioners and analogue to digital converter (ADC). The gear ratio of the worm gearbox is 1/15. The worm is made of case hardening steel and hardened to the value of 58\u201360 HRC. The worm gear wheel is made of bronze alloy. The worm is double threaded and the worm wheel has 30 teeth. The worm gearbox has been driven by a 2.2 kW, 3000 rpm three-phase motor and a variable speed controller has been used to drive the gearbox"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000216_tmag.2019.2940532-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000216_tmag.2019.2940532-Figure1-1.png",
        "caption": "Fig. 1. Configurations of the developed 6s/7p PE-HESFLM. (a) 2-D schematic. (b) 3-D view.",
        "texts": [
            " It should be mentioned that the proposed PE-HESFLM completely differs from the conventional double-sided machines, which are normally just a duplicate of the primary or the secondary to double the output power and thus still subject to space competitions within the primary [9]. This article is organized as follows. The configurations and operation principles of the developed machine are introduced first. Then, the optimization design is conducted. Finally, the electromagnetic performances of the optimized structures are studied comprehensively, followed by a conclusion in the end. The topology of the proposed 6-primary-slot/7-secondarypole (6s/7p) machine is shown in Fig. 1. As clearly shown, the PMs and field winding are placed into two different parts of the primary, with the top part of the primary accommodating the tooth-wound armature windings and PMs. On the other hand, the bottom part of the primary accommodates a toothwound field winding. It should be noted that the teeth of the 0018-9464 \u00a9 2019 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_14_0001040_j.addma.2019.06.003-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001040_j.addma.2019.06.003-Figure4-1.png",
        "caption": "Figure 4. DMA sample orientation. Angle was measure with respect to the loading direction.",
        "texts": [
            " It should be noted that due to variations in the diameter of the filament and imperfections in the build process, the final bead geometry is not a perfect rectangle, and the junction of contiguous beads still produces voids characteristic of FFF parts. Figure 3. Air gap application. Adapted from Rayegani et. al [10]. Dynamic Mechanical Analysis (DMA) tests were conducted using the tensile configuration of the RSA 3 (TA Instruments) DMA at ambient temperature. Three rectangular, single layer samples were produced with a thickness of 0.2 mm and dimensions of 35 mm by 6 mm. The samples were manufactured in three bead orientations: 0\u00b0, 45\u00b0, and 90\u00b0 where the angle is measured parallel to the load direction, as shown in Figure 4. Additionally, a sample of the parent material was compression molded with the same geometry and used for comparison. The parent material will be used to compare parts used in FFF applications against parts manufactured with other processes where the final product is a solid part, without induced anisotropy by the bead orientation. Subsequently, it is assumed that parts manufactured using these processes will have the same mechanical response as the compression molded sample. The same comparison was done for samples of the ultrasonic test"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002077_j.mechmachtheory.2020.104209-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002077_j.mechmachtheory.2020.104209-Figure3-1.png",
        "caption": "Fig. 3. Force diagram of the moving platform and i th limb.",
        "texts": [
            " Under the load effects, the reaction forces/moments generated from passive joints A i and B i can be measured in the frame A i -x i y i z i as f Ai = \u2212k L Ai x L Ai , m Ai = \u2212k A Ai x A Ai (7) f Bi = \u2212k L Bi x L Bi , m Bi = \u2212k A Bi x A Bi (8) The reaction forces F Ai / F Bi and reaction moments M Ai / M Bi generating from passive joint A i / B i measured in the frame O-xyz can be expressed as F Ai = O Ai f Ai , M Ai = O Ai m Ai (9) F Bi = O Ai f Bi , M Bi = O Ai m Bi (10) The force diagram of the moving platform is described as Fig. 3 . According to the Newton\u2019s 2nd law, the equilibrium equation of the moving platform can be formulated as \u2212 3 \u2211 i =1 F Ai + F P = 0 ;\u2212 3 \u2211 i =1 r A i \u00d7 F Ai \u2212 3 \u2211 i =1 M Ai + M P = 0 (11) where r Ai denotes the position vector of passive joint A i measured in the frame O-xyz . The structural compliance of limb body is further included into the kinetostatic model by modeling it as an elastic spatial beam with a number of discrete units e j ( j = 1, 2,... N 0 ) through the finite element (FE) method [42] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure25-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure25-1.png",
        "caption": "Fig. 25. Single-Bearing Traction Coupling and Friction Clutch for Clash Gearbox",
        "texts": [
            " This is often proposed but practically never carried out in practice, as it is a disagreeable solution from the point of view of length and cost ; also problems arise in connexion with whirling speeds and vibrations in the case of fast-running units. Furthermore, it seems wrong to resort to the use of a controlled clutch in series with a fluid coupling, which of itself is perfectly suited for taking up the drive. Perhaps the most satisfactory method of combining a fluid coupling and friction clutch is to use the single bearing coupling with a specially mounted clutch as shown by Fig. 25. The engine crankshaft carries the overhung weight of the impeller and casing only, since the runner with its shaft is supported by a spigot bearing in the boss of the impeller, and the primary half of the friction clutch mounted on this shaft is supported by a roller bearing in the nose of the gearbox, which is maintained in true alignment by a bell housing. The light spinning member of the clutch is mounted in a conventional way on the input shaft to the gearbox and supported by a spigot bearing in the driving plate of the clutch",
            " 98, it would be interesting to know whether this design introduced any difficulties in the form of fatigue of the material of the spring, which was repeatedly bending over two relatively small pulleys and one somewhat larger pulley; also how was the joint made at the ends of the spring ? On p. 106 reference was made to the use of a scoopcontrolled coupling for braking purposes. With this device the braking effect would apparently fall off progressively as the speed of the vehicle was reduced, and presumably the intention was to use it in combination with the normal brake, so as to reduce the wear on the latter. Fig. 25, p. 116, illustrated a combination of a traction coupling and a single-plate clutch. He understood that with this combination the plate clutch might be much smaller ; or in other words, the allowable pressure on the friction facing might be greater than with a normal friction clutch which had to take the full duty as the inherent flexibility of the fluid coupling relieved the momentary starting loads involved. Could the author indicate approximately (as a percentage) what extra pressure might be allowed on the facing of a single-plate clutch used in this manner as compared with one used without a fluid coupling ",
            " The coil spring pump was very similar to a coiled spring used for tachometer drives and did not involve fatigue since the pulley diameters were very large in relation to the section of the wire forming the coil. The joint was made by interlinking a simple loop formed in the two ends. The scoopcontrolled coupling, if used as a brake, was intended for use with a normal mechanical brake for stopping and holding purposes, as pointed out by Mr. Sykes. 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 182 PROBLEMS OF FLUID COUPLINGS Concerning Fig. 25, p. 116, he was not enthusiastic about the use of this construction for automobiles, and it was not really advisable to reduce the size of the friction surfaces as suggested by Mr. Sykes because one had to provide for the occasional driver who would persist in slipping his clutch in traffic, even though the fluid coupling rendered this unnecessary. It was best to apply the fluid coupling with a type of gearbox which eliminated any need for a clutch in series, for gear changing purposes. As regards the epicyclic creeping gear, Mr"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001250_j.triboint.2019.106028-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001250_j.triboint.2019.106028-Figure4-1.png",
        "caption": "Fig. 4. Change of the contact line length under different contact ratio.",
        "texts": [
            " The teeth surface contact lines of the double involute gears are inclined and intermittent, and the contact line first changes from short to long, then from long to short, until it disengages from engagement. According to the relationship between the transverse ratio \u03b5\u03b1 and overlap ratio \u03b5\u03b2 of the double involute gears, the length of the contact Fig. 1. Basic tooth profile of double involute gears. Fig. 2. Diagram of transverse engagement of double involute gears. Z. Yin et al. Tribology International xxx (xxxx) xxx line from entering the contact area to completely exiting the meshing area can be divided into the following two cases, as shown in Fig. 4. As shown in Fig. 5 (a), the distance between two adjacent contact lines is a transverse base pitch pbt , and the length of the contact line changes periodically with pbt . On the transverse of the gear, the moving distance of the i-th pair of contact teeth after the elapsed time t from the beginning is shown as follows: si \u00bc rb1\u03c91\u00bdmod\u00f0t;Tc\u00de\u00fe \u00f0i 1\u00deTc\ufffd \u00f0i\u00bc 1; 2;\u22ef; ceil\u00f0\u03b5\u00de\u00de (1) where Tc is the meshing period, Tc \u00bc 60=z1n1. i is the serial number of the meshing tooth, ceil\u00f0\u03b5\u00de function rounds \u03b5 to the nearest (higher) integer value, \u03b5 is the total contact ratio, \u03b5 \u00bc \u03b5\u03b1\u00fe \u03b5\u03b2, and its calculation method can be found in Ref"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003839_1.366700-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003839_1.366700-Figure1-1.png",
        "caption": "FIG. 1. Experimental setup.",
        "texts": [
            " Mixed cells were constructed by spincoating a PVCN solution onto the quartz substrate to a thickness of approximately 0.1\u20130.5 mm and baking at 70 \u00b0C for 1 h. The glass substrate was coated with DuPont polyimide 2555 ~PI! and then rubbed. This polyimide provides a small pretilt angle uPI'2 \u2013 4\u00b0 and a strong anchoring 8/83(1)/50/6/$15.00 \u00a9 1998 American Institute of Physics icense or copyright; see http://jap.aip.org/about/rights_and_permissions potential for the LC. The cell was filled with 5CB in the isotropic phase at ;50 \u00b0C, placed in a magnet field of approximately 8 kG which was at an angle a ~Fig. 1! to the cell normal, and cooled to room temperature under continuous exposure to polarized UV illumination from a mercury vapor lamp; the lamp\u2019s intensity I'0.1 mW cm22 ~Fig. 1!. Symmetric cells were constructed as above except that both the quartz and the glass substrates were treated with PVCN and neither substrate coated with polyimide or rubbed. The pretilt angle in the symmetric cells was measured by the \u2018\u2018magnetic null\u2019\u2019 method. The cell was rotated between the poles of the electromagnet until an orientation was found at which the phase retardation of the light passing through the cell did not depend on the strength of the applied magnetic field. This corresponds to a director orientation along the magnetic field, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002414_tmrb.2020.3011841-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002414_tmrb.2020.3011841-Figure2-1.png",
        "caption": "Fig. 2. Visualization of the relationship between inputs (\u03b8thigh, \u03b8\u0307thigh, and \u03b8\u0308thigh on the three co-ordinate axes) and outputs (A: ankle angle, B: ankle moment, C: knee angle, and D: knee moment, whose values are represented by using a spectral color map) from the complete recorded data of subject 1. A random forest model attempts to learn this relation from a subset of the complete dataset (training dataset) and then predicts on the remaining subset (validation dataset). The percentage gait cycle is also shown.",
        "texts": [
            " The predicted and expected output values for each gait cycle was resampled to 100 time samples (one sample each representing one percent of gait cycle). Further, the RMSE of prediction was calculated for each gait cycle percentage (each of the 100 samples) across all the validation trials from all the cross-validation iterations. Fig. 1 shows the time series measurements of the three inputs of subject 1 to the random forest model. It is the data for one gait cycle (time between consecutive heel contacts of right leg) as the subject performed level ground walking at self selected normal speed (mean speed = 1.28 m/s). Fig. 2 shows the input-output relationship for subject 1 which the random forest model tries to learn in case II. There is a homogeneous composition of the output values for different phases of a gait cycle with respect to the three-dimensional input values across different trials, making it easier for a random forest regression model to define the input-output relationships. Fig. 3 shows the R2 values and RMSE of the random forest predictions of the \u03b8ankle, \u03c4ankle, \u03b8knee, and \u03c4knee for both input combinations",
            " Although, the additional input (\u03b8\u0308thigh) did not provide visible improvement for \u03b8ankle and \u03c4ankle predictions, it enhanced the prediction accuracy for \u03b8knee and \u03c4knee prominently for at least two subjects. This would mean that the thigh angular acceleration, \u03b8\u0308thigh might not contain substantially surplus information helpful to the random forest model for inferring \u03b8ankle and \u03c4ankle during walking, but could be useful in predicting \u03b8knee and \u03c4knee. To further analyse the reason for this, we visualized the relation between the inputs to the random forest and each output to be predicted for subject 1 in figure 2. In a twodimensional input space (where inputs are \u03b8thigh, \u03b8\u0307thigh, case I), the input values overlap during the start and end of the gait cycle. This can be visualized for subject 1 in the thigh angle - thigh angular velocity plane (top view) of the three-dimensional plots in fig. 2. As a result, the regression model is expected to produce different output values (different colors in the spectral map) for similar input values (same \u03b8thigh \u2013 \u03b8\u0307thigh positions in the input space), which becomes infeasible. In this scenario, adding a third dimension (\u03b8\u0308thigh, case II) to the input space might help the random forest differentiate such output values, which were otherwise not discernible in a two-dimensional input space. This is true for \u03b8knee and \u03c4knee predictions where the Authorized licensed use limited to: Auckland University of Technology",
            " Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. third input dimension adds to the discriminatory power of the random forest. However, for \u03b8ankle and \u03c4ankle prediction, the third input seems to provide little additional information for discriminating the output values. For \u03c4ankle, this could be because the output values during the start and end of gait cycle are similar (similar colors during start and end of gait cycle in fig. 2B) and thus a two-dimensional input space is sufficient (please note that the \u03c4ankle prediction errors are low for case I for subject 1 during start and end of gait phase, fig 4). For \u03b8ankle predictions, the additional information provided by the third input, \u03b8\u0308thigh, might not be sufficient in discriminating the output values (note that only a small decrease in \u03b8ankle error was obtained in case II for subject 1 during the terminal gait phases, fig 4). In general, the prediction accuracy of the random forest method without any normalization of input features were comparable to the previous studies which employed machine learning algorithms for the prediction of gait variables with high accuracy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001546_s12206-020-0209-1-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001546_s12206-020-0209-1-Figure1-1.png",
        "caption": "Fig. 1. Bucket force diagram of wheel loader shoveling operation.",
        "texts": [
            " A distance appears between the bucket resistances when the bucket is turned over. Many factors affect the resistance, such as the type of material, the shape of the bucket, and the depth of the bucket insertion. This study mainly considers the influence of material type on the operation resistance. The other variables are fixed during the shovel loading test. Thus, the shape and insert depth of the bucket are not much of a concern. The working resistance is calculated according to different kinds of materials. Fig. 1 shows the force diagram of the bucket. (1) Inserting force: When the bucket is inserted into the material pile, inserting force is the reaction force of the material pile on the bucket. It includes shovel front cutting mouth and two side wall cutting mouth resistance forces, the friction resistance between the inner surface of bucket bottom and side wall and material, and the friction resistance of the outside surface of the bucket were in contact with the material. The inserting force of the bucket ( CP ) is calculated as follows [19]: 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003680_1999-01-0404-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003680_1999-01-0404-Figure7-1.png",
        "caption": "Figure 7. 2DOF vehicle model presentation",
        "texts": [
            " Several different complexity vehicle models have been developed for each platform. 2DOF MATRIXx, Full Car MATRIXx, and complete ADAMS mathematical models. While all these models have been used for different purposes, in the next section we present a simple 2DOF vehicle model. 2DOF VEHICLE MODEL \u2013 2DOF vehicle model is described with the Yaw dynamics equation and Side slip equation. In addition Pacejka tire model is used to calculate lateral forces. Yaw dynamics for 2DOF (bicycle) model [9],[10] Fig. 7, is described is follows: Jz dr/dt = a*Fyf - b*Fyr (37) where: Fyf Front tire lateral force Fyr Rear tire lateral force a CG to front axle distance b CG to rear axle distance Jz Yaw moment of inertia Side slip equation: M*Ve*(r +d\u03b2/dt) = Fyf + Fyr (38) where: M Vehicle mass Ve Vehicle velocity r Yaw rate \u03b2 Side slip angle Slip angle is calculated as: \u03b1f = \u03b2 + a*r/Ve - \u03b4 (39) \u03b1r = \u03b2 + b*r/Ve (40) where: \u03b4 is steering angle Cf Lateral force vs slip angle, front tire Cr Lateral force vs slip angle, rear tire"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002756_lra.2021.3061997-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002756_lra.2021.3061997-Figure1-1.png",
        "caption": "Fig. 1. Generation of circular field (dark red) and CF force (green) for an octagonal obstacle (light red) in 2D.",
        "texts": [],
        "surrounding_texts": [
            "In order to guide the robot to its goal pose we extend the definition of the attractor dynamics in the classical potential field approach with the proposed Velocity Limiting Controller (VLC) from [12]. For this, we first need to define an artificial desired velocity from the current robot position x \u2208 R3 and its goal positionxg \u2208 R3 in the form x\u0307d = kp kv (xg \u2212 x), where kp is the position gain and kv the velocity gain. The virtual forceFVLC is then calculated from the difference of the current robot velocity x\u0307 and the artificial desired velocity x\u0307d with FVLC = \u2212kv (x\u0307\u2212 \u03bdx\u0307d), where the factor \u03bd leads to the limitation of the velocity magnitude and is defined as \u03bd = min(1, vmax(x\u0307 T d x\u0307d) \u2212 1 2 ). The resulting control law is better suited for longer distances between robot and goal pose since the generated virtual force vanishes when the robot travels with the maximum velocity in the direction of the goal pose. This leads to a constant velocity magnitude except during acceleration, deceleration and in the vicinity of obstacles when the robot is subject to further virtual forces. We use the same approach for the orientation of the robot by computing an artificial desired angular velocity from the orientation errorxg,r \u2212 xr by using the quaternion difference of the current orientation q of the robot and the goal orientation qg as proposed in [28] with xg,r \u2212 xr = q0qg \u2212 qg,0q \u2212 q \u00d7 qg , where q0 and qg,0 are the scalar parts of the quaternions describing the current orientation and the goal orientation of the robot. Please note that the orientation is not used for the collision avoidance."
        ]
    },
    {
        "image_filename": "designv11_14_0002070_s10846-020-01269-y-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002070_s10846-020-01269-y-Figure2-1.png",
        "caption": "Fig. 2 Collision zone, repulsive zone, formation zone, and sensing zone around a spacecraft",
        "texts": [
            " It is assumed that there is always at least one spacecraft connected to the leader, namely, Nl \u2260\u2205. The virtual leader adjacency matrix is denoted by H = diag {h1, h2,\u2026, hn}, where hi = 1 denotes that the i-th spacecraft has the information of virtual leader, otherwise hi = 0. It is assumed that there is always at least one spacecraft with non-zero hi, namely, \u2211 n i\u00bc1 hi > 0. Then, the modified Laplacian matrix is defined as, eL \u00bc L\u00fe H \u00bc l11 \u00fe h1 \u22ef l1n \u22ef \u22f1 \u22ef ln1 \u22ef lnn \u00fe hn 24 35: \u00f013\u00de If the undirected graph G is connected, then all the eigenvalues of matrix eL are positive [56]. Figure 2 illustrates different zones around a spacecraft agent, considering inter-collision avoidance, obstacle dodging, and connectivity preservation, namely, collision zone, repulsive zone, formation zone, and a sensing zone, with the radius rCol, rRep, pdij, and r Sen, respectively. Collision with an obstacle or another spacecraft is assumed to occur when the distance between the two objects is less than rCol. Obviously, the radius rCol is larger than the sum of the radii of two adjacent objects. Further, it is assumed that rCol is smaller than the minimum desired relative distance between spacecraft agents"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002471_j.engfailanal.2020.104811-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002471_j.engfailanal.2020.104811-Figure10-1.png",
        "caption": "Fig. 10. FE model of the CDBG: (a) The 2-D FE model of the CDBG, which consist of 9504 2-D linear shell elements; (b) The 3-D FE model of the CDBG, which consist of 130,832 3-D linear solid elements. It should be noted that the spring elements represented for the support are not illustrated in this figure.",
        "texts": [
            " The thickness and area density of the inner ring is set to h1, 1, which is the thickness and area density of inner ring of the gear; then an equivalent thickness h2 and equivalent area density 2 are set for the outer ring as its stiffness and mass are influenced by the gear teeth. The shaft of the CDBG can also be approximated as a simplified shell structure to reduce the size of the FE model. In this model, the bearings to support the shaft are replaced by two sets of springs that are perpendicular to each other. The 2-D FE model established for the CDBG is illustrated in Fig. 10 (a). To determine the parameters h2 and 2 of this dynamic model, a 3-D FE model of the CDBG is also established to be treated as a standard model (shown in Fig. 10 (b)). The natural frequencies of these two FE models are calculated, then h2 and 2 are adjusted until the natural frequencies of the 2-D FE model are matched to the 3-D FE model. The final simulation results of are recorded in Table 3, which obviouly proves the accuracy of the 2-D FE model. In order to consider the contact stiffness of the teeth meshing with the CDBG, a spring with a stiffness of kc is applied at the contact point of the rotor-gear system. Because the meshing stiffness between the gear teeth could be changed by tooth error [28], flexibility of gear [29] and a fatigue crack occurred near the root of the tooth [9], the parameter kc is time-varying [30], which will significantly increase the complexity of the dynamic analysis of the CDBG",
            " The excitation frequency of gear-teeth meshing Hh can be expressed as the product of the operational speed of the CDBG g and the number of the gear teeth ng: =H n \u00b7h g g (4) Because the frequency of gear-teeth meshing is quite high and the meshing time is short (less than 1 ms), it is necessary to consider the high-frequency excitation as an impact load. For some more practical reasons, treat these high-frequency excitations as impact load can also reduce time steps significantly in the transient dynamic analysis of the CDBG. Therefore, the high-frequency excitation load applied at the CDBG can be approximated as an impact function (Fig. 10(a)), whereas the low-frequency excitation load can be represented by a sine function (Fig. 12(b)). The amplitude of the lateral vibration of the gas generator rotor is significant due to its high operational speed and weak front bearing stiffness. The rotor vibration will therefore stimulate an excitation on the CDBG. Since the contact stiffness of the gear pair in an actual gear-rotor system has strong nonlinearity, the excitation frequency applied to the CDBG by the vibrating rotor will contain multiple frequencies, which can be expressed as =H n \u00b7 ,l v r (5) where r is the operational speed of the gas generator rotor and =n 1, 2, 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003680_1999-01-0404-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003680_1999-01-0404-Figure5-1.png",
        "caption": "Figure 5. Gear valve assembly and schematics",
        "texts": [
            " The lumped parameter and distributed parameters line dynamic models are developed for the case that line dynamics frequency should be attenuated for the particular range of interest. Here we represent just the results of the correlated line model with simulated and tested frequency response, Fig. 4. 3 Simplified version of the line dynamic represented as a volume, is used first. The pump volume and line volume are integrated into a single volume. In the case that lumped or distributed parameters are used, the continuity equation is formulated differently. GEAR VALVE MODEL \u2013 Conventional gear valve structure is represented in Fig. 5. The gear valve model consist of orifice equations for all valve orifices in wheatstone bridge configuration. Torsion bar is acting as a torque sensor and is twisted by Steering wheel column torque on one side, and rack load on the other side. The twist angle determines the valve orifices opening and areas. Pinion shaft is coupled with the rack and supported in radial direction with the yoke. The gear valve model [9], is formulated as described bellow, and correlated to the test data obtained on the test bench Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001671_j.compstruct.2020.112468-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001671_j.compstruct.2020.112468-Figure3-1.png",
        "caption": "Fig. 3. Geometry of T-joint specimens and strain gauge locations.",
        "texts": [
            " Carbon UD fabric G0827 BB 1040 HP03 1F from M/s Hexcel Composites and EPOLAM 2063 epoxy resin system from M/s Axson Technologies are the main constituent materials used. Single cured ply thickness for these materials is 0.17 mm. Automated robotic machine having RS 522 tufting head with NM230 tufting needle (TN2.3) from M/s PFAFF Industriesysteme und Maschinen GmbH, KSL Germany [38] and TKT30 Kevlar thread from M/s Threads India Ltd is used for tufting the joints. T\u2010joint consists of stringer and rib, with mouse\u2010hole in the rib web, which are co\u2010cured along with skin panels. The geometry of T\u2010joint specimens experimented is shown in Fig. 3. Six T\u2010joint specimens are cut from one long cured segment. A single long segment is manufactured to avoid significant manufacturing variations between specimens. After curing, individual T\u2010joints are cut and prepared as per required dimension. Lay\u2010up sequence for individual elements, viz., skin, rib web and stringer along with thicknesses are given in Table 1. The individual elements are laid\u2010up on respective tool and preformed. Preforming is carried out at 80\u2070C under full vacuum in an air circulated oven",
            " T\u2010joints are tested by applying tensile load on the web while skin is clamped to the test fixture (Fig. 5). A specially designed pin loaded self\u2010aligning test fixture is used to avoid any bending caused on the specimen during testing. Quasi\u2010static loading is carried out at constant crosshead displacement speed of 1 mm/min in a universal testing machine. Two strain gauges are bonded to rib flange\u2010stringer flange cross\u2010over region and one strain gauge bonded to stringer web at the mouse\u2010hole region (Ref. Fig. 3). These locations are chosen in order to identify onset of failure in the specimen during loading. DIC technique is used to capture the full field strains and out of plane deformations of the skin. Digital cameras are placed below the specimen to record the images of skin during loading (Fig. 4). The bottom surface of the skin is randomly speckled for purposes of DIC measurements and the support fixture is modified to aid viewing of the speckled skin. The bottom surface of the skin is illuminated through LED lamps"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002352_j.mechmachtheory.2020.103945-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002352_j.mechmachtheory.2020.103945-Figure10-1.png",
        "caption": "Fig. 10. Definitions of virtual central link and virtual end link: (a) virtual central link, and (b) virtual end link.",
        "texts": [
            " In the coupling chain of a deployable unit, the moving link always passing through the SP perpendicularly is called the central link , and a pair whose axis is in the SP or which connects with the central link is called a central pair . From Table 1 , if the external loop linkage of a two-layer and two-loop linkage unit totally has even R pairs, its coupling chain must totally have odd R pairs. In this case, the unit has no central link and only has one central pair, whose axis is in the SP. To express output angular velocity of the coupling chain, a virtual central link parallel to the end link is assumed to fix to the central pair, shown in Fig. 10 (a). Please cite this article as: W.-a. Cao, Z. Jing and H. Ding, A general method for kinematics analysis of two-layer and twoloop deployable linkages with coupling chains, Mechanism and Machine Theory, https://doi.org/10.1016/j.mechmachtheory. 2020.103945 W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx 7 Otherwise, if the external loop linkage of a unit totally has odd R pairs, its coupling chain must totally have even R pairs. In this case, the unit has no end link, and has only one end pair , whose axis is in the SP. To express output angular velocity of the external loop linkage, a virtual end link parallel to the central link is assumed to fix to the end pair, shown in Fig. 10 (b). A fixed frame 0: o 0 -x 0 y 0 z 0 is attached to the geometrical center o of the fixed base, with z 0 axis along the normal of the base plane, x 0 axis along the normal of the symmetrical plane of the unit, and y 0 axis determined by the right hand rule. A moving frame e: o e - x e y e z e is attached to the geometrical center o e of the end link or the end pair in the SP, with x e axis along the normal of the symmetrical plane, z e axis along the projection of the axis of the end pair onto the SP, and y e axis determined by the right hand rule",
            "-a. Cao, Z. Jing and H. Ding, A general method for kinematics analysis of two-layer and twoloop deployable linkages with coupling chains, Mechanism and Machine Theory, https://doi.org/10.1016/j.mechmachtheory. 2020.103945 8 W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx Since a two-layer and two-loop deployable unit has a plane-symmetrical structure and has symmetrical motion, only chains 1, 2 and 3 of the unit in kinematics analysis need to be considered, shown in Fig. 10 . Based on the D-H parameter method [36] , from chains 1 and 2 of the unit, the homogeneous transformation matrix of the moving frame e with respect to fixed frame 0 can be expressed as 0 e T ( \u03b8l1 , \u00b7 \u00b7 \u00b7 , \u03b8l p , \u03b8u 1 , \u00b7 \u00b7 \u00b7 , \u03b8uq ) = 0 1 T ( \u03b8l1 ) \u00b7 \u00b7 \u00b7 p\u22121 p T ( \u03b8l p ) p p+1 T ( \u03b8u 1 ) \u00b7 \u00b7 \u00b7 p+ q \u22121 p+ q T ( \u03b8uq ) p+ q e T = \u23a1 \u23a2 \u23a3 n xe o xe a xe r xe n ye o ye a ye r ye n ze o ze a ze r ze 0 0 0 1 \u23a4 \u23a5 \u23a6 (7) where q denotes the number of R pairs of chain 1, p denotes the number of R pairs of chain 2, and \u03b8 l 1,\u2026, \u03b8 lp , \u03b8u 1, \u2026, \u03b8uq are joint angles of R pairs of chains 1 and 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure1-1.png",
        "caption": "Fig. 1. Geometrical diagram of 3-R(RRR)R+R HAM.",
        "texts": [
            " with intersecting axes. Thus, we have the 3-R(RRR)R parallel manipulator. The 3-R(RRR)R parallel manipulator consists of a fixed platform, a moving platform and three R(RRR)R chains. To adapt to the requirement of non-circular polarization antenna phase invariance, the polarization mechanism is connected in series between the 3-R(RRR)R parallel manipulator and the antenna reflector, so that the antenna reflector can rotate around the central axis of the moving platform. The novel 3-R(RRR)R+R HAM has 3R1T 4-DOF. Figure 1 shows the geometrical diagram of 3-R(RRR)R+R HAM. 3-R(RRR)R+R HAM is shown in Fig. 2. The fixed and moving platforms are expressed as P1 P2 P3 and Q1 Q2 Q3, respectively, and the branch chains are expressed as Pi Ji Gi Ki Qi (i = 1, 2, 3). The fixed coordinate system O \u2212 Ox Oy Oz and the moving coordinate system C \u2212 CxCyCz are, https://doi.org/10.1017/S0263574720001290 Downloaded from https://www.cambridge.org/core. University of Toledo, on 03 Jun 2021 at 19:25:40, subject to the Cambridge Core terms of use, available at https://www"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000292_s11432-019-1470-x-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000292_s11432-019-1470-x-Figure3-1.png",
        "caption": "Figure 3 (Color online) Representation of coordinate systems in the vertical plane.",
        "texts": [
            " The specifics of the hardware are given in Table 1. The developed UGSR prototype is composed of two telescopic modules and four rotate modules. Their parameters are listed in Table 1. This section derives the dynamic model of gliding motion in the vertical plane. The model is simplified by the following assumptions: (1) the robot is submerged in water and neutrally buoyant; (2) the robot does not move in the lateral direction; (3) the joint angles of all rotate joints are zero. The coordinate systems are defined using the right-hand rule in Figure 3. The world coordinate system, represented by Oxyz, is located at a fixed point, so the initial displacement of the robot is zero. The Ox and Oz axes lie in the forward direction and the gravitational acceleration direction, respectively. The body coordinate system, denoted by Obxbybzb, is initially located in the geometric center of the body. Its position coordinates with respect to Oxyz along the Ox and Oz axes are expressed as x and z, respectively. Obxb is aligned with the longitudinal axis of the body and points to the head module, whereas Obzb is perpendicular to Obxb and points downward in the vertical plane. The origin of the velocity coordinate system Ovxvyvzv coincides with the origin of Obxbybzb. The Ovxv axis aligns along the direction of the velocity with magnitude V , and Ovzv is normal to the Ovxv axis. The angle between Ox and Obxb is the pitch angle \u03b8, and the angle between Ovxv and Obxb is the angle of attack \u03b1. Counterclockwise pitch and attack angles are regarded as positive. In Figure 3, \u03b8 is negative and \u03b1 is positive. The gliding path angle \u03b3 defines the angle between Ox and Ovxv. The three angles are related by \u03b8 = \u03b3 + \u03b1. The kinematic model is then easily obtained as x\u0307 = V cos\u03b3, z\u0307 = \u2212V sin\u03b3. The robot consists of 6 modules, numbered j = 1, 2, . . . , 6 from head to tail. The elongations of telescopic modules 2 and 5 are represented by \u03b42 and \u03b45, respectively. The mass and volume of module j are expressed as mj and qj , respectively. Assuming uniform density of each module, the center of mass locates at the centroid, represented by rj with respect to frame Obxbybzb",
            " The buoyancy is denoted by B = \u2212mg, where m varies with \u03b42 or \u03b45. As the rg and rb are functions of the elongations \u03b42 or \u03b45, moments are exerted by gravity and buoyancy if \u03b42 6= \u03b45. These moments are described as Tg = \u2211 mjgrgcos\u03b8 \u2212mhgrhsin\u03b8, Tb = \u2212mgrbcos\u03b8. Hydrodynamic forces and moments are generated when the body moves through the fluid. The force is resolved into its drag D and lift L components, which are parallel and perpendicular to the velocity direction in the vertical plane, respectively (see Figure 3). The pitching moment of the couple is indicated by MDL. Based on aircraft dynamics, the hydrodynamics of the drag, lift and moment are modeled in terms of the variables \u03b1 and V D = (CD0 + CD\u03b12)V 2, L = (CL0 + CL\u03b1)V 2, MDL = (CM0 + CM\u03b1+ Cq\u03c92)V 2, where CD0, CD, CL0, CL, CM0, CM and Cq are hydrodynamic coefficients obtained by a computational fluid dynamics simulation, and \u03c92 is the angular velocity of \u03b8. The dynamic model of gliding motion for the UGSR is given by V\u0307 = 1 M1 (\u2212D \u2212 gsin\u03b3m0), \u03b3\u0307 = 1 M1V (L\u2212 gcos\u03b3m0), \u03b1\u0307 = \u03c92 \u2212 1 M1V (L\u2212 gcos\u03b3m0), \u03c9\u03072 = 1 J2 (MDL + Tg + Tb), (1) where the mass M1 along the Obxb axis is assumed equal to the mass along Obzb"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002690_tbme.2021.3053374-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002690_tbme.2021.3053374-Figure2-1.png",
        "caption": "Fig. 2. (a) Meshed solid model of the composite energy-storage-andreturn (ESAR) prosthetic foot for a finite element analysis. Mode shapes and strain distributions of the composite ESAR prosthetic foot (thin black line is undeformed model): (b) first mode, (c) second mode, (d) third mode, and (e) fourth mode.",
        "texts": [
            " The strain distributions were numerically evaluated using FEM to identify the ideal placement of the piezoelectric patches on the prosthetic foot for experimental testing of the physical prosthetic foot. Previous research has shown that the damping effectiveness of small patches is maximized when the patches are placed at the locations of the highest bending moments [29]. The meshed solid model of the composite ESAR prosthetic foot was constructed by commercial FEM software, ANSYS (Canonsburg, Pennsylvania) (Fig. 2a). The model was meshed with 20-node brick elements and 5784 elements were generated (3D Solid 191 elements). As a boundary condition, the proximal end of the composite foot was clamped and thus all degrees of freedom at this end were constrained. The material properties of the foot (composite Young\u2019s modulus: 64\u00d7109 Pa, Poisson\u2019s ratio: 0.16, and mass density: 1600 kg/m3) were obtained by Fig. 1. Piezoelectric shunt damping test setup of the composite ESAR prosthetic foot (SPR-5-S-N, \u00d6ssur). Note that positions A and B indicate the locations for the piezo patches (dimensions: 5",
            " 4): 40 Hz (first mode), 112 Hz (second mode), 408 Hz (third mode), and 556 Hz (fourth mode). The analytically determined modal frequencies from the FEM analysis (Table I), with the experimental values added as a reference, indicate that the first three analytically determined modal frequencies are well matched with the experimental frequencies, allowing identification of high strain locations on the foot from the FEM analysis and placement of the piezoelectric patches. The mode shapes and strain distributions of the composite ESAR prosthetic foot are shown in Fig. 2b-e. In the first mode, the main deformation was the bending motion of the composite prosthetic foot, and large displacement occurred at the free end. At the clamped area, the strain was maximized and the strain at the free end was minimized. The second mode was similar to the buckling shape of the composite prosthetic foot. Thus, moderate strain occurred around the heel as well as the clamped area. In the third mode, most strain occurred in the area between the free end and the heel. In the fourth mode, similar levels of strains occurred in four regions; around the free end, the heel, above the heel, and the clamped poximal end"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001208_j.triboint.2019.105999-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001208_j.triboint.2019.105999-Figure7-1.png",
        "caption": "Fig. 7. EVD sealing device.",
        "texts": [
            " The device has an internal Nomenclature D rod diameter, m PRMS1 contact pressure RMS of external stroke, MPa E elastic modulus, GPa FN, F\u2019N contact force of seal pair, MPa PAmin minimum contact pressure of external stroke static seal, MPa Ff1, Ff2 axial friction of seal pair, MPa fRMS friction force RMS in the external stroke, MPa PARMS contact pressure RMS of external stroke static seal, MPa H displacement of piston rod external stroke, m PA contact pressure of LA, MPa Sx(f)max frictional power spectrum peak ho oil film thickness, mm V surface speed of rod, m/s havg average film thickness, mm V0 leakage of external stroke, ml hRMS oil film thickness RMS, mm WA maximum pressure gradient L contact length of sealing area, mm WE wavelet packet energy entropy LA the point with the highest pressure gradient \u03c5 Poisson\u2019s ratio \u03b2 the lowest energy ratio P fluid pressure, MPa \u03c1 density of the liquid,kg/m3 P(x) contact pressure of seal pair, MPa \u03b7 dynamic viscosity,Pa* s P0 compensating pressure, MPa \u03c4\u00f0x\u00de friction stress of sealing pair, MPa P1 pressure of the test cylinder, MPa \u03c4  max maximum friction stress, MPa Pmax maximum contact pressure of seal pair, MPa Fig. 1. Elastomer combination seal. Fig. 2. Spring energized seal ring. X. Zhao et al. Tribology International 142 (2020) 105999 pressure cavity in the seal elastic ring, can uses an external supercharging device to inject high-pressure oil into the cavity to increase the sealing contact pressure and restore the sealing function of the seal [8], as shown in Fig. 7. Although the EVD system solves the problem of seal compensation pressure regulation, there are still some problems that need to be studied. For example, in order to reduce the heat and wear of the seal contact area due to the increase in the seal contact pressure, how to adjust the compensation pressure to achieve a balance between the contact force and the friction is one of the main challenges. In order to solve this problem, a design scheme of pressure compensation piston rod seal is proposed in this study, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002731_j.mechmachtheory.2021.104285-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002731_j.mechmachtheory.2021.104285-Figure4-1.png",
        "caption": "Fig. 4. Schematic of the tooth slice and cantilever beam model.",
        "texts": [
            " The manufacturing error can be expressed as follows [26] : e m ( t ) = 0 . 5 F p sin ( 2 \u03c0\u03c9 s t + \u03c8 s ) + 0 . 5 f \u2032 i sin ( 2 \u03c0\u03c9 m t + \u03d5 m ) (58) where F p is the total cumulative pitch deviation, f i \u2019 is the tangential deviation of single tooth, \u03c9 s and \u03c8 s are the gear rotation frequency and its initial phase angle respectively, \u03c9 m and \u03d5m are the meshing frequency and its initial phase angle respectively. Therefore, the similarity ratio of F p and f i \u2019 can be obtained, \u03bbe m = \u03bb f \u2032 = \u03bbF p (59) The method shown in Fig. 4 can be used to calculate the mesh stiffness of helical gear. The gear is divided into N segments along the width of the teeth. When the number of segments is enough, each tooth segment can be equivalent to a spur gear. The mesh stiffness of each spur gear can be calculated based on potential energy method [ 27 , 28 ]. Under the action of force F , the bending, shear and compression energy of the tooth can be expressed as { U b = F 2 / ( 2 K b ) U s = F 2 / ( 2 K s ) U a = F 2 / ( 2 K a ) (60) where K b ,K s , and K a are equivalent stiffness corresponding to bending, shearing and compression deformation respectively. The spur gear is equivalent to the cantilever beam model shown in Fig. 4 . According to the beam theory, the potential energy of the tooth under bending, shearing and radial compression deformation can be represented by \u23a7 \u23aa \u23a8 \u23aa \u23a9 U b = \u222b d 0 M 2 2 E I x dx U s = \u222b d 0 1 . 2 F 2 b 2 G A x dx U a = \u222b d 0 F 2 a 2 E A x dx (61) where F b , F a and M can be obtained from Eq. (62) . { F b = F cos \u03b11 F a = F sin \u03b11 M = F b x \u2212 F a h (62) Combined with Yang\u2019s research results of Hertz contact stiffness [29] , the tooth bending stiffness K b , shear stiffness K s , radial compression stiffness K a and Hertz contact stiffness K h can be obtained as follows: \u23a7 \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23a9 1 K b = \u222b d 0 ( x cos \u03b11 \u2212h sin \u03b11 ) 2 2 EI x dx 1 K s = \u222b d 0 1 . 2 cos 2 \u03b11 GA x dx 1 K a = \u222b d 0 sin 2 \u03b11 EA x dx 1 K h = 4 ( 1 \u2212v 2 ) \u03c0EW (63) In Eq. (61) - (63) , I x and a x represent the area moment of inertia and area of the section where the distance between the section and action point of meshing force is x , and the tooth width is W , the other parameters are shown in Fig. 4 . The deformation of the gear body also has a great influence on the mesh stiffness. The deformation of the gear body can be expressed as Eq. (64) [ 28 , 30 ]. \u03b4 f = F cos 2 \u03b1m W E ( L \u2217 ( u f S f )2 + M \u2217 ( u f S f ) + P \u2217 ( 1 + Q \u2217tan 2 \u03b1m )) (64) The coefficients L \u2217, M \u2217, P \u2217, Q \u2217 can be approached by polynomial functions X i \u2217(h f i , \u03b8 f ) = A i \u03b82 f + B i h 2 f i + C i h f i \u03b8 f + D i \u03b8 f + E i h f i + F i (65) the values of A i ,B i ,C i ,D i ,E i , and F i are given by reference [30] . The parameters u f ,s f ,h fi = r f / r int and \u03b8 f are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001775_tim.2020.3006682-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001775_tim.2020.3006682-Figure7-1.png",
        "caption": "Fig. 7. Final anti-buckling case design.",
        "texts": [
            " The geometry of the anti-buckling case was then defined accordingly (see \u00a7 IV-B) taking into account the necessity of allowing the entrance of sheets having various thickness, and the requirements regarding the magnetic measurement tools. Further calculations were realized to ensure the mechanical stability of the system. A conservative value of the calculated contact pressure was chosen (\u03c3contact = 0.1 MPa) and applied uniformly on the case surface. This allowed defining the material in which it should be realized, remembering that it must be magnetically and electrically neutral. The anti-buckling case can be seen in Figure 7: it consists in a two-parts case of PEEK (Poly-Ether Ether Ketone) material, conferring to the system an excellent geometrical stability, high shear modulus, low friction. Moreover, the material is non-magnetic and electrically insulating. The geometry allows the insertion of sheet having different thicknesses thanks to an adjustable clearance, resulting in a \u201dsandwiched\u201d laminated sample. Once the sheet is in its position, a set of sixteen (16) nylon screws are used to fix the system setup. A calculation (Kellerman-Klein formula) was realized in order to define the torque necessary to be applied on each screw in order to compensate the contact pressure of the buckling plate without over-stressing the sheet in its thickness. Indeed, this stress could alter the material response to the magneto-mechanical excitation, which is to be avoided. Some features were added in order to conveniently place up to four (4) H-coils \u2013 or other thin local probes, such as ARM or Hall effect sensors \u2013 at various distances from the sheet surface. In Figure 7, two of them are visible as cuts on the case side (red arrows). Around the middle of the case, the secondary coil is wound, enveloping the H-coils. A space dedicated to the positioning of strain gauge was also foreseen. In order to apply mechanical loads of various intensities, a universal tension/compression test stand is used. This machine allows the generation of uniaxial loadings up to 2500 N, which corresponds to stresses up to 50 MPa for a laminated sheet of cross-section S=100x0.5 mm2. A commercial tester was chosen to realize the mechanical excitation: the MECMESIN R\u00a9 MultiTest-d 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001041_s40998-019-00214-6-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001041_s40998-019-00214-6-Figure2-1.png",
        "caption": "Fig. 2 The forces acting on the Quadrotor",
        "texts": [
            " However, this method will be hard to implement in practice because adding one input variable will greatly increase the number of control rules (as the example in the work of Fu et\u00a0al. (2014)); the constructing of fuzzy control rules is an even more difficult task and it needs more computation efforts. Hence it is better to design a fuzzy controller that possesses the fine characteristics of the PID controller by using only the error e(t) and the change rate of error e\u0307(t) as its inputs for the lateral and longitudinal motion x and y (Fig.\u00a02). The PID-type FC has nonlinear behavior that strongly depends on the values of the scaling factors Ke , Kde , Kpd and Kpi , where the authors in Qiao and Mizumoto (1996) have realized a linear structure of fuzzy controllers, in which, their outputs can be written as: According to Qiao and Mizumoto (1996) the fuzzy controller (in Eq.\u00a026) can behave approximately like time-varying PD controller named PD-type FLC , where A, P and D are (26)fcpd = f (e, de, t) = Kpd(A + Pkee + Dkdee\u0307) (27) fcpi = Kpi \u222b Of dt = kpi \u222b (A + Pkee + Dkdee\u0307)dt = kpiAt + kpikdeDe + kpikeP\u222b edt the fuzzy controller components calculated after linearization of the functionf(e,\u00a0de,\u00a0t) [a detailed demonstration is presented in Qiao and Mizumoto (1996)]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure12-1.png",
        "caption": "Fig. 12 Schematic diagram of working relief angle",
        "texts": [
            " The changing trend of the cutting force of cutter B is contrary to the working rake angle, which proves that the cutting force has negative correlation with the working rake angle. Therefore, the curved rake face can not only ensure that the working rake angle always stays in a reasonable range but also reduce the fluctuation amplitude of cutting force to make the cutting process more stable. Reasonable working relief angle is the premise of avoiding interference between the flank face and the machined tooth surface. As shown in Fig. 12, the working relief angle \u03b1o appears as the intersection angle between two straight lines in plane Po. One is the intersection line of plane Po and plane Ps, and the other one is the tangent line of the intersection curve of plane Po and the flank face. N2 and t\u03b1 represent the direction vectors of them respectively. Since N1, N2, and t\u03b1 are coplanar, t\u03b1 can be expressed by: t\u03b1 \u00bc sin\u03b10N1 \u00fe cos\u03b10N2 \u00bc cos\u03b10N1\u2212cos\u03b10ve21 \u00f019\u00de At the same time, t\u03b1 locates in the tangent plane of the flank face, so t\u03b1\u2219nh \u00bc sin\u03b10N1\u2219nh \u00fe cos\u03b10N2\u2219nh \u00bc sin\u03b10N1\u2219nh\u2212cos\u03b10ve21\u2219nh \u00bc 0 \u00f020\u00de where nh represents the normal vector of the flank face which can be obtained by Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002218_j.mechmachtheory.2019.103771-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002218_j.mechmachtheory.2019.103771-Figure11-1.png",
        "caption": "Fig. 11. Finite element analysis model of the lifting mechanism.",
        "texts": [
            " It is worth noting that node Q is a prismatic joint whose two adjacent units belong to different components; thereby two different elastic angles should be set at such kind of nodes. Since the displacements at node A in both directions are bounded by the frame assumed to be rigid, only one elastic angle is set. The rod IQ is regarded as the cantilever beam to eliminate the freedom of the rigid-body, so that the whole mechanism can satisfy the \u201cinstantaneous structure hypothesis\u201d [27] . Consequently, there is no generalized coordinate set at node I , and the finite-element model of the lifting mechanism is finally presented in Fig. 11 . In the lifting mechanism, 15 generalized coordinates are used for the motion analysis, which are numbered by the system numbers. The generalized coordinate U L in the global coordinates using system numbers is denoted as U L = [ U L1 U L2 \u00b7 \u00b7 \u00b7 U L15 ]T (13) Correspondingly, vector U\u0308 Lr consisting of the rigid-body acceleration for each node in the lifting mechanism is defined as U\u0308 Lr = [ U\u0308 L1r U\u0308 L2r \u00b7 \u00b7 \u00b7 U\u0308 L15r ]T (14) According to the rigid-body dynamic results of the forging manipulator, Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000567_j.jsv.2016.04.020-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000567_j.jsv.2016.04.020-Figure2-1.png",
        "caption": "Fig. 2. (a) Structure of a ballscrew drive and (b) the schematic model of the screw-nut interface.",
        "texts": [
            " In Section 2, after introducing the model of multi-subsystem connected via a wedge mechanism, the receptance coupling equation was derived and briefly discussed. In Section 3, the modeling approach for ballscrew drives based on the derived receptance coupling was illustrated and verified. In Section 4, a sensitivity-based position-dependent dynamics analysis method was given, after the modal characteristics analysis. Finally, conclusions were drawn in Section 5. A ballscrew-nut joint connects the ballscrew and the sliding carriage, marked as component I and component II in Fig. 2 (a), and transforms the ballscrew's rotation into sliding carriage's axial motion. Due to the flexibilities of the components, the joint also carries complicated transmission and transformation of the vibrations. These vibrations mainly include the rotational, axial, and bending vibrations of the ballscrew being a continuous beam, and the axial, lateral, pitching, yawing, rolling, and even higher modes of vibration of the sliding component being a multi-rigid-body system. The dynamic couplings among these vibrations at the ballscrew-nut joint are realized by imposing restriction on the interface coordinates of the vibration subsystems. Six DOF interface coordinates are considered for each component and the restriction are mathematically abstracted as that of a screw-nut-couple and a cylinder-bush-couple, as illustrated in Fig. 2(b). A three-dimension five-DOF ballscrew-nut joint model is presented in Fig. 2(b). In the model, the interface is divided into three parts: a flexible screw-nut couple and two flexible cylinder-bush in XY and YZ plane. For simplicity, the screw-nut couple is unwound circumferentially, and becomes a planar-wedge couple. The interface condition presented by the schematic model was given in Eq. (1). That determines the dynamic couplings among above mentioned vibrations, and will be used in the receptance coupling modeling and analysis of the ballscrew drives. kjn\u00f0\u03b4Iy cos \u03b2\u00fed\u03b8Iry sin \u03b2\u00de \u00f0\u03b4IIy cos \u03b2\u00fed\u03b8IIry sin \u03b2\u00de \u00bc f jn kjx\u00f0\u03b4Ix \u03b4IIx \u00de \u00bc f jx kjrz\u00f0\u03b8Irz \u03b8IIrz\u00de \u00bc Tjrz ( kjz\u00f0\u03b4Iz \u03b4IIz \u00de \u00bc f jz kjrx\u00f0\u03b8Irx \u03b8IIrx\u00de \u00bc Tjrx ( (1) where, \u03b4Ix; \u03b4 I y; \u03b4 I z; \u03b8 I rx; \u03b8 I ry; \u03b8 I rz are the six linear and angular displacements of the ballscrew's coordinate where the ballscrew connected to the nut; \u03b4IIx ; \u03b4 II y ; \u03b4 II z ; \u03b8 II rx; \u03b8 II ry; \u03b8 II rz are the six linear and angular displacements of the nut assumed rigid; d is the kinematic diameter of the ballscrew; \u03b2 is the helical angle of the ballscrew which is identical to the wedge angle; kjn is the normal stiffness which is orthogonal to the wedge interface due to the screw-ball-nut compliances, and all the compliances are assumed to come from the rolling ball's contact point with the screw and the nut [11]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002037_s00170-020-06366-8-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002037_s00170-020-06366-8-Figure7-1.png",
        "caption": "Fig. 7 \u201cMachining jig\u201d MSHAM design concept",
        "texts": [
            " These four substrate blocks, 25 mm in thickness, had the following characteristics and features: \u2022 Serve as the substrate for the additive build \u2022 Overlap as much shape as possible to reduce print time \u2022 Function as a machining jig for downstream finishing operations \u2022 Contain screw holes for mounting onto the build plate and CNC machines \u2022 Contain location dowel pin holes for precise position alignment on the build plate and in CNC finishing operations As the four inserts were to be printed in one single build, the thickness of all base substrates had to be the same. Figure 6a and b show the \u201creduce print time\u201d design for Wedge Insert \u201cA\u201d with the hybrid-build concept and the premachined base substrate, respectively. The majority of the base substrate will become an integral part of the insert after the printing operation. Figure 7 shows a different MSHAM design concept in which the majority of the substrate block will serve as machining jig only. However, a small amount of the substrate material will end up being the cavity shape. The total print volume with this setup was 483,303 mm3. Once the four base substrates were prepared by CNC machining, they were mounted onto the build plate ready for printing (Fig. 8). After printing, all four inserts were simply removed by unscrewing them from the build plate, and they were ready for finish machining",
            " Further tempering of the fabricated mould insert by conventional T6 or hot isostatic pressing (HIP) heat treatment could increase its ductility with a small reduction in tensile strength [21, 22]. It is worth noting from the test results that the choice of the base substrate material and the positioning of the substrate block in the modified mould insert design for hybrid-built is essential. In this study, the cavity shape of the Cavity Base Insert was designed to be part of the substrate block (Figs. 7 and 9). The powder-substrate interface is only a couple of millimetres from the CCC (Fig. 7). However, based on the result of the tensile test, the substrate block should be located on the base side of the mould insert so that the cavity shape, together with CCC, could be produced from AlSi10Mg, the stronger of the two materials. One may argue that if a higher strength material could be used for the base substrate, the position of the interface would be of less importance. However, a strong fusion bond between the power and the substrate is still the most critical factor in ensuring success in the MSHAM method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001966_j.mechmachtheory.2020.104136-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001966_j.mechmachtheory.2020.104136-Figure7-1.png",
        "caption": "Fig. 7. (a) Gears meshing. (b) Time-varying mesh stiffness.",
        "texts": [
            " In the gear box, gear set works as an amplifier of the torque. In recent literature on driveline modeling, the gear sets are assumed as rigid elements and the gear box is considered as a lump element, whose inertia is calculated based on the gear ratio. In fact, when the gear set transmits torque by teeth meshing, the number of meshing teeth varies with the rotation speed of the gears, which further results in varying contact stiffness of the gears. The gear set meshing and the contact stiffness are shown in Fig. 7 . With the assumption that the bearings, which support the gear shaft are rigid, and that the axial force of the helical gear set is not considered, the gear set model can be simplified to have only torsional motion issues. Fig. 8 presents the diagram of the gear set motion and the equivalent torsional motion model. The dynamic model of the gear set can be obtained based on the Newton\u2019s second law, and can be expressed by Eq. (8) . { J i \u0308\u03b8i + F d r i \u2212 T i = 0 J k \u0308\u03b8k \u2212 F d r k + T k = 0 F d = k mik ( t ) d is ( t ) + c mik ( t ) d is ( \u02d9 t ) dis ( t ) = r i \u03b8i \u2212 r k \u03b8k \u2212 e ik dis ( \u02d9 t ) = r i \u02d9 \u03b8i \u2212 r k \u02d9 \u03b8k \u2212 \u02d9 eik (8) Where, r i, r k is the base circle radius of the gear i and k, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002814_j.ijmecsci.2021.106405-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002814_j.ijmecsci.2021.106405-Figure11-1.png",
        "caption": "Fig. 11. Beam model with the specific loading case a) ; formation of a plastic hinge on the beam under load b) and s tress-strain evolution in the critical cross-section c) .",
        "texts": [
            " Geometrical features: the geometrical dimension of the unit cell a) ; definition of the initial gap b) ; formulation of the geometrical model c) ; deflection of the ligament d) \ud835\udc63 \ud835\udc63 b d f i a i u \ud835\udf00 y t i h b \ud835\udf00 \ud835\udc40 \ud835\udf00 \ud835\udc40 i \ud835\udc39 t \ud835\udc39 a \ud835\udc39 s d \ud835\udc51 1 c \ud835\udc51 5 m c t Furthermore, the deflection of the ligament is determined: = \ud835\udc42\ud835\udc36 \u2212 \ud835\udc42\ud835\udc37 ; \ud835\udc63 = \ud835\udf0c \u2212 \ud835\udc42\ud835\udc37 (10) The following equation is obtained: = \ud835\udf0c \u2212 \ud835\udf0c \u22c5 cos \ud835\udf03 (11) And by replacing Eq. (8) in Eq. (11) the center angle \ud835\udf03 is computed y numerically solving the following equation: 2 \u22c5 \ud835\udf03 \u22c5 \ud835\udc63 \ud835\udc3f \ud835\udc5a = 1 \u2212 cos \ud835\udf03 (12) Once a value for the center angle \ud835\udf03 was obtained, the curvature ra- ius \ud835\udf0c can be computed using Eq. (8) The model of the beam (ligament) under the effect of the loading orce F is presented in Fig. 11 . Considering the model presented in Fig. 11 a) , the applied force F s producing a bending moment M that is loading the ligament under pure bending load case. With the load force increase, the stresses are ncreasing, and a plastic hinge is developed, as presented in Fig. 11 c) . The strain produced on the outer fiber (on the gap side) is estimated sing: = \ud835\udc61 2 \u22c5 \ud835\udf0c (13) The strain should be more than a reference value corresponding to ield stress ( \ud835\udf0ey ) to develop a plastic hinge. Using the constitutive law of he steel as elastic-perfectly plastic [ 45 , 68 ], the stress-strain evolution n the cross-section is presented in Fig. 11 c) . Two areas characterize the cross-section. The elastic region with a eight defined by t \u2032 and a plastic region corresponding to the are defined y t \u2032 /2 and t /2. The strain required to reach the yield point is given by: \ud835\udc66 = \ud835\udc61 \u2032 2 \u22c5 \ud835\udf0c\ud835\udc66 (14) The moment loading the beam is defined by [ 42 , 45 ]: = \ud835\udc40 \ud835\udc52,\ud835\udc5d = \ud835\udf0e\ud835\udc66 \u22c5 \ud835\udc61 \u20322 \u22c5\ud835\udc64 6 + 2 \u22c5 \ud835\udf0e\ud835\udc66 \u22c5 [ \ud835\udc64 \u22c5 ( \ud835\udc61 2 \u2212 \ud835\udc61 \u2032 2 ) \u22c5 1 2 \u22c5 ( \ud835\udc61 2 + \ud835\udc61 \u2032 2 ) ] (15) Using the following relations: = \ud835\udc61 2 \u22c5 \ud835\udf0c ; \ud835\udf00 \ud835\udc66 = \ud835\udc61 \u2032 2 \u22c5 \ud835\udf0c\ud835\udc66 (16) Eq. (15) becomes: = \ud835\udf0e\ud835\udc66 \u22c5 \ud835\udc61 2 \u22c5\ud835\udc64 4 \u22c5 [ 1 \u2212 1 3 ( \ud835\udc61 \u2032 \ud835\udc61 ) 2 ] (17) The vertical load F is exerted over three ligaments (one out of four s the loading ligament); thus, the force is given by: = 3 \u22c5\ud835\udc40 \ud835\udc45 \ud835\udc56 + \ud835\udc61 \u22152 (18) The force developed on the node is (there is one active ligament and hree are transmitting): \ud835\udc62 = 2 \u22c5 \ud835\udc39 (19) Considering the numbers of cells ( n H ) along the circumference, the xial force is: \ud835\udc4e = \ud835\udc5b \ud835\udc3b \u22c5 \ud835\udc39 \ud835\udc62 (20) The gaps, the number of cells along the vertical direction, and the hape of the ends provide the solution of the plateau length ( d p ) before ensification: \ud835\udc5d = ( \ud835\udc5b \ud835\udc49 \u2212 1 ) \u22c5 \ud835\udeff (21) In some cases, it was observed that the plateau length extends by /2 \u2022 \ud835\udeff allowing the deformation of the ligaments close to the ending ircular rings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003626_piae_proc_1922_017_035_02-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003626_piae_proc_1922_017_035_02-Figure3-1.png",
        "caption": "FIG. 3.",
        "texts": [
            " It shows that fo r any ordinary road speed, speed is the governing factor, and for practical purposes we may say that the pressure on the r o d varies m the square of the speed, a result confirmed by expensive experiments in America. The equation (1) may be 1-gaided as a term of a Fourier series, and tlius the foregoing conclusions apply to any form of road irregularity subject to the interpretation of wave length. The effects of speed may be shown with much generality and clearness by simple veotor methods, see Fig. 3 . The momentum M of the being W we have F I G . 1. F I G . 2. vehicle increases by dM in the time dt. For a given aaguhr change, dt k inversely as v, and since M varies as v the rate of change of momentum varies as u2, i.e., the reaction on any given road varies as t h square of the speed. Thus, the road reaction is proportional to the weight, and by looking for the conditions under which road reaction is zero, we have the well-known effect of bounoe or leaving the road. The vertical acceleration upwards is given by (2) and the reaction R by (3)",
            " 465 is to increme the virtual wave length of the noad obstacle. practical results 80 far apparent me, therefore:The 1. With heavy wheels the speeds must be limited. 2. Road irregularities are aggravated and not smoothed out by 3. Increased wheel diameter smooths out the road, and, but for fast traffic. increased mass, would always be good. This follows from consideration of bounce. PART IV.-S~MPLE SPRUNG SYSTEM OF TWO MASSES-THE SPRING CART. Suppose, now, that instead of a single mass, we have two masses ml and .nu, c o ~ e ~ t e d by a spring (see Fig. 3), the mass of which may be ignored or allocated in part to each of the masse,s ml and m2. Let s be the stiffness of the spring in lb. per foot. and let XI and x2 be the heights of m, and m2 measured from their equilibrium positions when they are at rest on the level part of the road. Then we have the Newtonian equation of motion of m, . . . . . . . . . . . . . . . . . . . . . . . . ( 7 ) m,x, + 8 (x2 - zI) = The conclusions to be drawn from this equation are: - 1. The motion of the upper mass is governed by (a) The ratio s/mz",
            " The physical significance of equation ( 7 ) is that the motion of a car body is independent of the sprung maas of the car provided that the ratio of spring stiffness to sprung mass is constant. From this Dr. Reissner has argued that a light car might be as comfortable as a heavy car. This would be correct i f a car could be truly ROWELL. 30 at The University of Auckland Library on June 5, 2016pau.sagepub.comDownloaded from 466 'I'HE INSTITUTION OF' AIITOMORILE ENGINEERS. represented so simply as in Fig. 3. As a matter of faat, the treatment of motor-car suspensions as two masses aonn'wtd by a sprin,g is only a crude approximation, which, while s'ommhat better than our first method wheaein the 'car was treated as a singlc particle, is very far from being correct. * (8) *@a) Re-arranging (7) we have:........................... z2 + sx,/m, = sxl/?nl' XI = y = (hi21 (1 - cos U t 2 9 7 / X ) z., + sx,/na, = (sh/2rr~,) (1 - cos c f 2 r / ~ ) . or assumiqg a path for the axle as before, Le., assuming . "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000556_0142331216645179-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000556_0142331216645179-Figure1-1.png",
        "caption": "Figure 1. Conceptual design of the Turac.",
        "texts": [
            " Also, especially for UAVs used in urban areas, it is desirable to have a high cruise speed, payload capacity, hover flight and vertical-takeoff-and-landing (VTOL) capability. To combine all these features in a single design, Turac has been developed. It has hover-to-transition and transition-tohover capabilities by using tilt rotors and a coaxial main lifting fan. The body is designed as a blended wing which increases aerodynamic performance in flight. Also, interchangeable wings with different spans provide adequate lift force for different missions (Ozdemir et al., 2014). Important conceptual specifications of the Turac are shown in Figure 1. One-half and one-third scale prototypes are manufactured and flight tests are performed (Aktas et al., 2014, 2015) as shown in Figures 2 and 3. Six-DoF equations of motion of the transition flight are obtained and the aerodynamic effects are evaluated (Vuruskan et al., 2014; Yuksek et al., 2016). Flight control systems have an important role in ensuring the success of UAV missions. When the controlled system becomes more complicated, it is necessary to design a controller which is robust against faults and structural failures"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003659_mssp.1998.0190-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003659_mssp.1998.0190-Figure4-1.png",
        "caption": "Figure 4. The best selection of measurement points when n\"3; con\"guration composed of measurement points is a regular triangle.",
        "texts": [
            " The denominator and the numer- ator of equation (17) can be written as equations (18) and (19), respectively, J 1 \"D d 12 ]d 23 D2\"1 4 D d 12 D2 D d 23 D2 sin2/ (18) J 2 \"D d 12 D2#D d 13 D2#D d 23 D2\"D d 12 D2#D d 12 #d 23 D2#D d 23 D2 (19) where / is the angle between d 12 and d 23 . Assuming that the perimeter of the triangle composed of the three measurement points is constant, the condition satisfying Max [J 1 ] and Min [J 2 ] can be easily derived /\"603 and D d 12 D\"D d 23 D\"D d 13 D . (20) Indeed, this con\"guration is a regular triangle as shown in Fig. 4. 5. SELECTION OF EXCITATION POINTS AND DIRECTIONS [R r ] and [R q ] are closely related to the excitation conditions. In this section, \"rst the determinants of [R r ] and [R q ] are exactly expressed with angular acceleration vectors. Second, the condition numbers of these matrices are approximated by the physical quantities, and the worst excitation conditions are derived analytically. Finally, some guidelines for the selection of excitation points are suggested. 5.1. MATRIX CHARACTERISTICS OF [R r ] If the angular acceleration vector h j is de\"ned as h j ,aK j i#bG j j#cK j k, j\"1, 2, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002856_s13369-021-05654-z-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002856_s13369-021-05654-z-Figure3-1.png",
        "caption": "Fig. 3 a Schematic of the building approach (side view) (upskin, core and downskin); b Schematic of the tracks of laser interactions and the distance between each track is called Hatch distance (top view); c schematic of the orientation of the laser tracks w.r.t the previous layer is at 67\u00b0 for core printing strategy [6]",
        "texts": [
            " Each traverse of the laser beam melted and further solidified the material layer by layer to create the plate. Moreover, the beam offset was taken to compensate for the dimensional error, in which the laser beam was shifted by half of the curing diameter of the spot size. The process parameters used for building the AM plates are given in Table\u00a02. During the printing of each plate, the layer thickness of each layer was kept at 30\u00a0\u03bcm. The plate was built with the three-layer system, i.e., upskin, core and downskin, and is shown schematically in Fig.\u00a03a. The two layers first printed at the bottom are called \u2018downskin\u2019; the three layers printed in the end are called \u2018upskin\u2019 and the rest of the layers between upskin and downskin are called \u2018core\u2019 of the AM 1 3 plate. Each layer following another in the regular sequence of the core was rotated by 67\u00b0 during printing, as shown in Fig.\u00a03c. The CAD representation of the printing orientation of DMLS plates is shown in Fig.\u00a04a, followed by the actual printed plates on the substrate base plate as shown in Fig.\u00a04b. Eight plates of additively manufactured aluminum alloy (AlSi10Mg) of size (100 \u00d7 50 \u00d7 3) mm were selected for post-processing using shot blasting. Each plate was further sectioned into four specimens of size (50 \u00d7 25 \u00d7 3) mm (Fig.\u00a01a) using WEDM (Wire-cut electric discharge machining). The sectioned specimen plates were cleaned in acetone to remove any debris or loose particles present on the surfaces before subjecting them to post-processing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002048_0142331220966427-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002048_0142331220966427-Figure3-1.png",
        "caption": "Figure 3. The propeller group coordinate system.",
        "texts": [
            " Section 3 introduces the ESO-based omnidirectional control and control allocation strategy. Section 4 analyzes the observer stability and cloesd-loop stability. Extensive numerical simulations are presented in Section 5, while conclusions and future directions will be explained in Section 6. Let FW : OW : XW ,YW ,ZWf g denote the world inertial frame fixed on the ground, and FB : OB : XB,YB,ZBf g denote the body frame attached to the vehicle body at the center of gravity, as seen in Figure 2. There are also four frames of the propeller groups FPi : OPi : XPi ,YPi ,ZPi f g shown in Figure 3. These frames rotate around their XPi axes, w.r.t the body frame. As usual, the rotation matrix WRB stands for the orientation of the vehicle body frame w.r.t world inertial frame while BRPi stands for the orientation of the i-th propeller group frame w.r.t body frame. By denoting the propeller tilting angle around the axis XPi with ai, it follows that BRPi =RZ p 2 i+ p 4 RX (ai), i= 1 . . . 4: \u00f01\u00de Also let OB Pi =RZ p 2 i+ p 4 l 0 0 2 4 3 5, \u00f02\u00de be the origin of FPi in FB with l being the distance between OPi and OB"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.38-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.38-1.png",
        "caption": "Fig. 10.38 The structure of the cylindrical chopper",
        "texts": [
            " The drive power of the model is in the range of 50\u201355 kW. Compared with the cone-shaped chopper, it has the advantages of compact equipment, small site area, small motor power, long residence time, high stability, and so on. However, the wear of blade is fast, and the change of blade is very complicated. In the case of eight sets of 50-model cylinder chopper, the knife is made of No. 20 chromium molybdenum steel and carburized, which can be only used for the production about 300 tons of NC. The structure of the cylindrical chopper is shown in Fig. 10.38. The diameter of the rolling cutter is 280\u2013400 mm, which is equipped with 38\u201350 pieces of steel knives with the blade thickness of 4\u20136 mm. The roll knife has a certain angle with the axis (such as 10\u00b0\u20138\u00b0), which usually operates at high speed, 900\u20131000 r/min. There are four sets of stator knives located evenly around the roll knife, in which the blade is parallel to the axis. Compressed air or high-pressure water is employed for adjusting the knife. The feeding inlet of the chopper is equipped with a feeding impeller, by which the NC slurry is continuously fed evenly between the rolling cutter and the bed knife"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000180_chicc.2019.8865370-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000180_chicc.2019.8865370-Figure1-1.png",
        "caption": "Fig. 1: Reference coordinate frames",
        "texts": [
            " And the conclusions are drawn in the last Section. Quadrotor helicopter is a highly nonlinear underactuated system, which has six degree of freedom and four control inputs. It is assumed that the quadrotor is a rigid body and the structure is completely symmetrical. For the convenience of describing the quadrotor helicopter dynamic model, an inertial reference coordinate frame Oxyz relative to the earth and a body reference coordinate frame Oxbybzb are defined. The typical structure and these two reference coordinate frames are shown in Fig.1. To describe the behavior of the quadrotor, the absolute position vector is expressed as \u03b6 = [x, y, z]T , body angular velocity vector is expressed as \u03b7 = [p, q, r]T and Euler angle vector is expressed as \u0398 = [\u03d5, \u03b8, \u03c8]T . Euler angles are respectively roll angle \u03d5, pitch angle \u03b8, and yaw angle \u03c8 with the assumption of \u2212\u03c0/2 < \u03d5 < \u03c0/2, \u2212\u03c0/2 < \u03b8 < \u03c0/2 and \u2212\u03c0 < \u03c8 < \u03c0. Since Euler angles are small during actual flight, we can assume \u0398\u0307 = \u03b7. The rotation matrix from the body coordinate reference frame to the inertial reference coordinate frame is given by R = \u23a1\u23a3 c(\u03c8)c(\u03b8) c(\u03c8)s(\u03c6)s(\u03b8)\u2212 c(\u03c6)s(\u03c8) ux c(\u03b8)s(\u03c8) c(\u03c6)c(\u03c8) + s(\u03c6)s(\u03c8)s(\u03b8) uy \u2212s(\u03b8) c(\u03b8)s(\u03c6) uz \u23a4\u23a6 (1) where the terms c("
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003023_s11071-021-06591-0-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003023_s11071-021-06591-0-Figure4-1.png",
        "caption": "Fig. 4 Robotic fish platform",
        "texts": [
            " 3, where the three base stations of the navigation and position system are located surrounding the pool. The mobile station is installed in the internal part of the fish body. Then, the real-time location, swimming direction and speed of the RF are determined via the analysis of the relative position between the mobile and base stations. The information of the actual speed and the yawing angle is sent and demonstrated in the upper computer wirelessly. In order to better fit the body wave of the biological fish, the physical prototype of the RF is depicted in Fig. 4. The fusiform head is used to accommodate the battery, the DSP-based master control and wireless communication modules, mobile station, and gyroscope. The flexible fish body with four connected joints covered by a latex skin and the caudal fin are also designed. The physical mass (M) is 4:2 kg, while the lengths of the RF and fusiform head (d in Fig. 1) are 0:67 m and 0:34 m, respectively. The RF body links consist of DC motors as well as other supporting components. For fitting the tail of the real fish, the aluminum alloy skeleton is used in the RF caudal fin to connect to the fourth joint motor by using silica gel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000203_j.sna.2019.111729-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000203_j.sna.2019.111729-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagrams of (a) Au nanorods/p-NIPAAm hydrogel composite in a macroporous silicon membrane below lower critical solution temperature (LCST), (Au nanorod/hydrogel composite: pink, silicon membrane: gray, silicon dioxide: yellow) (b) Operation of NIR light actuated macroporous silicon membrane filled with A h",
        "texts": [
            " u nanorods with a specific absorption band are incorporated in -NIPAAm hydrogel, and the resulting hydrogel composite infills he pores to form arrayed hydrogel columns. At the temperaure below the LCST of the hydrogel, light can be transmitted to he other end of the macroporous membrane via transparent pIPAAm hydrogel columns as waveguides. As the temperature of he hydrogel columns in/near the surface of the macroporous silcon membrane is raised over the LCST by external stimuli, the ransparency of p-NIPAAm hydrogel changes to opaque and light ransmission is blocked as shown in Fig. 1(b). Simultaneously, the ydrogel columns are deswelled and release liquid in hydrogel olumns through micron-sized pores. In the proposed design, the xternal stimulus is NIR light, which is corresponding to the SPR bsorption band of incorporated Au nanorods. sors and Actuators A 301 (2020) 111729 2.2. Materials N-isopropylacrylamide (NIPAAm) and Silver Nitrate (AgNO3) were purchased from TGI America. N,N\u2019-Methylene-bisacrylamide (MBAAm), ammonium persulfate (APS), ascorbic acid (C6H8O6), N,N,N\u2019,N\u2019\u2013 tetramethylethylenediamine (TEMED), and Polyvinylpyrrolidone (PVP) were purchased from VWR"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001485_j.mechmachtheory.2019.103776-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001485_j.mechmachtheory.2019.103776-Figure4-1.png",
        "caption": "Fig. 4. A modified Bennett linkage [11] . (a) Deployed configuration, (b) fully folded configuration, the red arrows x 1 , y 1 , z 1 denote the reference coordinate systems. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " (51) The corresponding self-stress wrench is given with one arbitrary constant \u03b2 as F = \u03b2 ( 0 \u22121 0 0 \u22121 0 )T , (52) this wrench represents a combination of forces and torques with the pitch h = 1, in the form of the wrench as in Eq. (16) . The quadratic form in Eq. (31) for this linkage is Q = \u03b2 ( 4 4 \u221a 3 \u2212 3 4 \u221a 3 \u2212 3 12 \u2212 3 \u221a 3 ) (53) which is positive definite if \u03b2 > 0, thus this RCRCR linkage is a first-order infinitesimal linkage. 6.3. Modified infinitesimally mobile Bennett linkage The modified Bennett linkage in Fig. 4 has two sets of link lengths, and it becomes a special Bennett linkage with finite mobility when all link lengths are equal. This modified Bennett linkage was thought to possess second-order infinitesimal mobility based on position constraint analysis in [11] but the method in this paper shows it only possesses first-order infinitesimal mobility. The D-H parameters of this modified Bennett linkage are a 41 = a 12 = a, a 23 = a 34 = b ; \u03b112 = \u03b134 = \u03b1, \u03b123 = \u03b141 = \u2212\u03b1; d i = 0 , i = 1 , \u00b7 \u00b7 \u00b7 , 4 . (54) The screw coordinates of joints in the fully folded configuration in Fig. 4 (b) are \u03be1 = ( 0 0 1 0 0 0 )T , \u03be2 = ( 0 \u2212 sin \u03b1 cos \u03b1 0 \u2212a cos \u03b1 \u2212a sin \u03b1 )T , \u03be3 = ( 0 0 1 0 \u2212a \u2212 b 0 )T , \u03be4 = ( 0 sin \u03b1 cos \u03b1 0 \u2212a cos \u03b1 a sin \u03b1 )T . (55) The first-order motions are one dimensional with the basis matrix as N = ( \u2212 2 b a + b cos \u03b1 1 \u2212 2 a a + b cos \u03b1 1 )T . (56) The self-stress wrenches are given with three arbitrary constants \u03b21 , \u03b22 , \u03b23 , F = \u239b \u239d 1 0 0 0 0 0 0 0 0 1 0 0 0 \u2212a 0 0 0 1 \u239e \u23a0 T ( \u03b21 \u03b22 \u03b23 )T . (57) It can be seen that only the coefficient \u03b21 is left after applying Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000931_s00170-019-03553-0-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000931_s00170-019-03553-0-Figure2-1.png",
        "caption": "Fig. 2 a Dimensions of the miniature tensile specimen in mm. b Specimen loaded in the self-aligning grips with extensometer set. c Specimen broken after the test",
        "texts": [
            " The deposited material was characterized using a Helios NanoLab 600 SEM equipped with an Oxford energy-dispersive X-ray spectrometer, a PANalytical X\u2019Pert research diffractometer, a Struers Duramin hardness tester, and an Instron universal testing machine. The analyses involved scanning electronmicroscopy, energy-dispersive X-ray spectroscopy (EDS), optical microscopy, X-ray diffraction (XRD), and tensile testing towards the characterization of the material properties and microstructures. The tensile characterization of the material was carried out by testing miniature specimens on an Instron universal testing machine. A Hansvedt DS-2 wire EDM was used to cut these specimens to the dimensions shown in Fig. 2. To ensure perfect uniaxial testing, the Instron system was customized with self-aligning grips (Fig. 2). In case of a grip-specimen misalignment, the grip reposition was minimally based on the need to avoid torquing or bending of the miniature tensile specimen. The tensile tests were carried out at a closed-loop controlled strain rate of 0.015/min till a strain of 1% (for accurate offset yield strength measurement) and then the rate was ramped up to 0.5/min till failure [19]. The initial stages of the study involved the fabrication and characterization of deposits made from blended powder feedstocks containing elemental nickel and elemental copper powders"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure1-1.png",
        "caption": "Fig. 1. Arbitrary rigid body rotating about axis passing through O.",
        "texts": [
            " While the mass and centre of mass location of a rigid body can be determined by static tests, the mass moment of inertia tensor (inertia) can only be measured through dynamic testing [6]. The general approach is to study the dynamic response of the system to initial conditions and/or forcing input to determine its inertial properties. An arbitrary rigid body rotating about a rotation axis passing through O has inertia properties that can be described by the tensor IO expressed in a body-fixed reference frame with its origin at O (Fig. 1). IO is described by: IO \u00bc IOxx IOxy IOxz IOxy IOyy IOyz IOxz IOyz IOzz 0 B@ 1 CA \u00f01\u00de The dynamics of the angular velocity vector x t\u00f0 \u00de \u00bc xx t\u00f0 \u00de;xy t\u00f0 \u00de;xz t\u00f0 \u00de T of the rigid body expressed in coordinates of its body-fixed frame are governed by the Newton-Euler equation s \u00bc IO _x t\u00f0 \u00de \u00fex t\u00f0 \u00de IO x t\u00f0 \u00de\u00f0 \u00de \u00f02\u00de where s is the net torque vector acting on the body about O, expressed in coordinates of its body-fixed frame. When the rotation is constrained to the vertical axis zO, Fig. 1, such that the angular velocity vector is expressed as x t\u00f0 \u00de \u00bc 0;0; _h t\u00f0 \u00de T \u00f03\u00de the equations of motion reduce to: sx sy sz 0 B@ 1 CA \u00bc IOxz\u20ach t\u00f0 \u00de IOyz _h t\u00f0 \u00de2 IOxz _h t\u00f0 \u00de2 \u00fe IOyz\u20ach t\u00f0 \u00de IOzz\u20ach t\u00f0 \u00de 0 B@ 1 CA \u00f04\u00de Eq. (4) has been traditionally used to experimentally determine IOzz; IOxz, and IOyz. Genta et al. [6] classified the experimental methods for estimating the moment of inertia, IOzz, into two main categories: angular acceleration methods and oscillation methods. In acceleration methods, an object is subjected to an external torque or initial condition and the dynamic response is observed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000556_0142331216645179-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000556_0142331216645179-Figure4-1.png",
        "caption": "Figure 4. Body and Earth coordinate systems.",
        "texts": [
            " In teh fourth section, a fault detection and isolation system is designed for the sensor fault scenario. In the fifth section, the virtual sensor is designed by using a reconfigurable Kalman filter to overcome sensor signal cut-off and to maintain system stability during post-fault flight. In the sixth section, the designed fault tolerant control system is integrated on the six-DoF nonlinear mathematical model and simulation studies are performed for the sensor cut-off fault scenario. Equations of motion (EoM) of the UAV are obtained by using Newton\u2019s second law. In Figure 4, the body and Earth axis systems are shown. Six-DoF motion of the UAV is defined by using nonlinear force, moment, orientation and position equations as shown in equations (1)\u2013(4) (Nelson, 1998). m( _U +QW RV )=FGx +FAx +FTx m( _V +RU PW )=FGy +FAy +FTy m( _W +PV QU)=FGz +FAz +FTz \u00f01\u00de _PIxx +RQ(Izz Iyy) ( _R+PQ)Ixz =LA +LT _QIyy RQ(Ixx Izz)+ (P2 R2)Ixz =MA +MT _RIzz +PQ(Iyy Ixx)+ (QR _P)Ixz =NA +NT \u00f02\u00de P= sin (u) _c+ _f Q= sin (f) cos (u) _c+ cos (f) _u R= cos (f) cos (u) _c sin (f) _u \u00f03\u00de at Middle East Technical Univ on May 17, 2016tim"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003688_1.1304914-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003688_1.1304914-Figure3-1.png",
        "caption": "Fig. 3 The Euler angles and reference frames used to parameterize the rotor\u2019s rotation relative to the track",
        "texts": [
            " We define a set of body-fixed vectors $et1 ,et2 ,et3% that form a right-handed orthonormal basis \u00abt which corotates with the track. Here, et1 and et2 lie in the plane that contains the track, while et3 is normal to the track\u2019s plane. Restricting the nutation angle u to be constant, the angular velocity vector of the circular track is vt5c\u0307E32c\u0307et3 (1) where a dot over a variable denotes differentiation with respect to time. Only two independent Euler angles are necessary to specify the current orientation of the rotor relative to the track, as viewed in Fig. 3. The first rotation, through an angle a, is about et3 , and causes the axle to rotate around the track. Following this rotation, it is convenient to define a right-handed orthonormal basis $e1 ,e2 ,e3% such that e1 is parallel to the axle and e3 is parallel to et3 . Thus, the vector e2 lies in the plane spanned by et1 and et2 . Note that this basis is neither corotational with the track nor corotational with the rotor. For the second rotation, the rotor spins about e1 through an angle g. Corotational with the rotor is the right-handed orthonormal basis \u00abr5$er1 ,er2 ,er3%",
            "org/ on 04/28/20 tion for the final independent Euler angle a and two equations for the constraint moment. Evaluating epi\u2022H\u03075epi\u2022Mr yields z~hv\u03071!2@v\u030731za\u0307v21~12h!v1v2#1 sz l2 v150, v\u030722za\u0307v31~h21 !v1v35 k2 l2 , (19) hv\u030711z@v\u030731za\u0307v21~12h!v1v2#1 s l2 v15 k3 l2 sin b . The derivation of these scalar equations uses the relationships in ~3! and ~9!, as well as a dimensionless parameter h: h5 l1 l2 . (20) Using ~16!, ~19!1 becomes an uncoupled, nonlinear ordinary differential for a(t). We simplify this governing equation by introducing a new angle ~see Fig. 3!: the phase angle d~ t !5a~ t !2c~ t !. (21) Additionally, we nondimensionalize by introducing t5c\u0307ot (22) where c\u0307o is the initial condition for the precession rate (c\u0307o 5c\u0307(0)). Substitution into ~19!1, yields d91ad81bc82 cos d2~bc91cc8!sin d2dc82 sin 2d 5ec92ac8. (23) Here, each apostrophe denotes differentiation with respect to the dimensionless variable t ~e.g., d85dd/dt and d95d2d/dt2! and the constant coefficients are a5 z2n 11z2h , b5 zh sin u 11z2h , c5 zn sin u 11z2h , d5 1 2 ~h21 !sin2 u 11z2h , e5 z2h1cos u 11z2h "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001834_j.addma.2020.101547-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001834_j.addma.2020.101547-Figure5-1.png",
        "caption": "Fig. 5. Polished substrate.",
        "texts": [
            " The deposition track of the disc-shaped specimen (2 r= 25.0mm) in the DED process is shown in Fig. 4: hatch spacing was w=2.0mm and there were 11 fusion lines in each layer. Figs. 4a and 4b show the deposition tracks of the first and second layers, respectively. As shown in Fig. 4c, the deposition orientations of the first and second layers were perpendicular to each other. The substrate used in this study is a thick substrate with a thickness of 16mm, which is made of 304-stainless steel. As shown in Fig. 5, the bottom of the substrate was polished. As shown in Fig. 6, a special fixture is designed in the experiment. The substrate was coaxially clamped on the center hole of the support frame by two bolts. The fixture consists of bolts, support frame and springs. The gap between the substrate and the fixture ensures that the substrate can move freely in the x- and y-directions. The substrate is only constrained by the spring force in the z-direction. Fig. 7 shows the entire experimental setup for in-situ monitoring using the CGS system in the experiment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002152_978-981-13-3549-5-Figure5.18-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002152_978-981-13-3549-5-Figure5.18-1.png",
        "caption": "Fig. 5.18 Interaction geometry and mechanics of a wheel on deformable terrain [96, 101]. a Interaction of wheel and deformable terrain and b force of a wheel moving on a simplified slope",
        "texts": [
            " 3 and 4, it remains to understand the wheel\u2013terrain interaction for implementation of obstacle avoidance through the geospatial modeling of the planetary exploration wheeled mobile robot. The wheel\u2013terrain interaction model contains two parts: geometrical contact and interaction mechanics, corresponding to the terrain geometry and mechanical properties, respectively. (1) Geometrical Contact Calculation: To simulate the interaction mechanics with high fidelity, the interaction area of a wheel should be calculated and considered as an important factor when moving on soft soil. Figure 5.18a shows the interaction area of a wheel moving on rough terrain. The DEM of the terrain is provided, and the known parameters are (xw, yw, zw) and \u03d5w. They are corresponding to the central position and the yaw angle of a wheel, respectively. The interaction area is simplified as an inclined plane determined by points P1, P2 and P3. The inclined plane P1P2P3 can be represented in At (x \u2212 x1) + Bt (y \u2212 y1) + Ct (z \u2212 z1) 0. (5.6) 5.4 The Future Research Priorities 133 Equations of predicting the coordinates of points P1, P2, P3 and At , Bt , Ct are presented in ref. [101]. The contact can be considered as a wheel with a simplified slope determined by points P1, P2, P3, as shown in Fig. 5.18b, where { e} and { w} are coordinate systems with the same orientation and different origins, at the end point and wheel center, respectively. The wheel sinkage is then determined by z r \u2212 |At (xw \u2212 x1) + Bt (yw \u2212 y1) + Ct (zw \u2212 z1)|\u221a A2 t + B2 t + C2 t . (5.7) The contact can be decomposed into climbing up a slope with angle \u03b8 cl and crossing a slope with angle \u03b8 cr: { \u03b8cl arcsin[(\u2212At \u2212 Bt tan \u03d5w)/X1] \u03b8cr arcsin[Ct (At tan \u03d5w \u2212 Bt )/X2] , (5.8) where X1 \u221a C2 t (1 + tan2 \u03d5w) + (At + Bt tan \u03d5w)2 and X2 \u221a X3[A2 t + C2 t + 2At Bt tan \u03d5w + (B2 t + C2 t ) tan2 \u03d5w]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002546_tie.2020.3039203-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002546_tie.2020.3039203-Figure5-1.png",
        "caption": "Fig. 5. The experimental platform of the PMSM drive with NCS.",
        "texts": [
            " Finally, the SMC approach adopts the continuous linearized form instead of the signum function for a boundary layer solution, and its gain parameter  should be selected considering the system sensitivity to variations and highfrequency noises. With the less information of the unmodeled high-frequency dynamics, the trial and error approach may be a considerable tuning method. Normally the higher gains, the better performance, the cost is that the system is more susceptible to the unmodeled high-frequency dynamics, e.g., measurement noises [24], [29]. Table I summarizes the tuning guideline of parameters in the proposed control approach. IV. EXPERIMENT VALIDATION A. Experiment set-up Fig. 5 shows the experimental platform, where one PMSM acts as driving motor and the other serves as load motor, and their shafts are connected via an elastic coupling. The angular displacements of the motor shafts are measured by the including rotary encoders. The servo drivers operate in torque control mode, receiving the torque targets as the control inputs and transmitting the positions as the feedback signals. The designed controllers are implemented in automation software (TwinCAT 3.0). The EtherCAT protocol is applied for digital data transmission"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002091_0954405420978039-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002091_0954405420978039-Figure5-1.png",
        "caption": "Figure 5. Tooth contact pressure distribution for the hypoid gear pair manufactured by usual applied machine tool settings.",
        "texts": [
            " The data for the thermal elastohydrodynamic calculation are presented in Table 2. The calculations were made for 31 instantaneous mating positions of the gear pair. The calculated pressure distributions along the potential contact lines of tooth surfaces for these 31 instantaneous mating positions and the calculated values of operating characteristics of the gear pair (maximum tooth contact pressure, pmax, angular displacement error of the driven gear, Df2 max, and efficiency of the gear pair, h) are presented in Figure 5, for the case when the machine tool settings are calculated by the usually applied method. The next figure, Figure 6, shows the tooth contact pressure distributions and the values of operating characteristics for the case when the pinion teeth are manufactured by the head-cutter of optimized geometry and by applying optimal variation in machine tool settings governed by equation (13). It can be considered that by the introduction of the optimized manufacture procedure the maximum tooth contact pressure is reduced from 1520MPa to 1181MPa, the maximum angular displacement error of the driven gear from 22"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002816_j.optlastec.2021.107024-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002816_j.optlastec.2021.107024-Figure3-1.png",
        "caption": "Fig. 3. Design of the benchmark part. Dimensions are in mm.",
        "texts": [
            " A statistical analysis based on the experimental results is presented to highlight the most influential parameters and to provide regression models. Measurements of displacement are repeated after annealing, to investigate those modifications induced by the heat treatment. In the following section the plan of experimental activities is presented (Section 2.1); laser parameters implemented are discussed (Section 2.2); the manufacturing stage (Section 2.3) and the experimental measurements are described (Sections 2.4 and 2.5). M. Mele et al. Optics and Laser Technology 140 (2021) 107024 Fig. 3 shows the benchmark part used for testing (support structures are not represented here for the sake of simplicity). The design concept of the chosen support structures is shown in Fig. 4. This linear geometry is extruded up to the part\u2019s surface. To ease their subsequent removal, all the supports are connected to the part by using teeth with height and width equal to 1.50 mm and 0.80 mm, respectively. All the dimensions are proportional to the distance dsl between the supports (Fig. 4). This design allows the effective removal of non-transformed powder before cutting all the supports",
            " Heat treatment is performed by raising the oven temperature from 20 \u25e6C to 1050 \u25e6C at 500 \u25e6C/min and maintaining this M. Mele et al. Optics and Laser Technology 140 (2021) 107024 temperature for 2 h. The cooling stage is carried out by switching off the oven and waiting for the chamber to cool down to room temperature before the part is extracted. Contact and non-contact measurements are implemented to evaluate possible distortions induced by the manufacturing process. The distance between the base of the benchmark and the lower surface of the overhanging portion of the specimen (Fig. 3) is considered. This nominal value of the aperture, hn, can be calculated such as in Eq. (1): hn = \u230a 2 + 5 + Loh \u00d7 sin(\u03b1oh) \u2212 2 \u00d7 tan(\u03b1oh) hL \u2309 \u00d7 hL (1) Due to the layer height hL, it is worth underlining that Eq. (1) takes into account the round value of the nominal dimension. The mechanical measurements of displacements are made by a Vernier calliper with accuracy \u00b10.05 mm. All the benchmarks are examined in ten different positions. The average value and the standard deviation are recorded and used for the following comparative analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001834_j.addma.2020.101547-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001834_j.addma.2020.101547-Figure8-1.png",
        "caption": "Fig. 8. Schematic of substrate: (a) experimental measurement area in cladding process, (b) the unclad area and clad area in cladding process.",
        "texts": [
            " In this study, we performed in-situ monitoring for measuring substrate deformation in the DED process using a CGS system to investigate the substrate deformation characteristics at different stages. These results include the relationship between the deformation evolution and the number of fusion lines in the cladding process of the first layer, and the relationship between the deformation evolution and the deposition track in the cladding process of the 6th fusion line in the first layer. For better clarification, the measurement area shown in Fig. 8 was divided into clad area and unclad area. And the specimen was divided into different parts during the first layer cladding process to describe the characteristics of substrate deformation at different stages is shown in Fig. 9. The interference fringes on the substrate at different stages during the first layer cladding process are shown in Fig. 10. From Eq. (4), we can obtain surface slope from the interference fringes. Further, the surface curvature can be obtained by differentiating the surface slope and the surface shape can be obtained by integrating the surface slopes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001312_tie.2019.2952780-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001312_tie.2019.2952780-Figure9-1.png",
        "caption": "Fig. 9. Prototype of the electromechanical actuator highlighting different sections such as the pulley-belt and clutch mechanisms.",
        "texts": [
            " From the FEA model, it is found that when 50 kN is applied, a deformation of 0.055 mm is produced in the drawbar. Moreover, based on the distribution of the safety factor and the considering the allowable stress of (material), a minimum safety factor of 6.62 can be selected for the drawbar. This safety factor is considered as the minimum safety factor for the whole actuator system during manufacturing process. Using the dimensions mentioned in Table II, a prototype of the electromechanical actuator is manufactured as shown in Fig. 9. The different important components of the manufactured electromechanical actuator such as the drive motor with pulley-belt mechanism, spindle system housing, clutch mechanism are highlighted separately in Fig. 9. It is noteworthy that during any machine operation, a wide constant power-speed range must be ensured to deal with various machining conditions. For an example, a high torque and low speed is desired when cutting a hard-to-machine material and synchronoized tapping, whereas high power and high speed must be ensured to cut aluminum alloys [16]. Moreover, when the two modes of machining are considered individually, during clamping mode, the electromechanical actuator should maintain a high clamping force to hold the work piece perfectly under a low speed condition, Thus the drive motor should ensure a high torque to produce the required clamping force",
            " To analyse the performance of the electromechanical actuator under the clamping operation, the thrust at the drawbar and the chucking force produced by the chuck jaws attached at the end of the drawbar are tested in this subsection. 1) Test for the thrust produced by the drawbar during the clamping: To analyse the thrust produced by the drawbar during the clamping mode, the drive motor shaft is coupled with the motor side shaft of the electromechanical actuator using a pulley-belt mechanism as shown in Fig. 9 and the motor is run at the low speed winding mode. The pulley-belt mechanism used for the test bench has a pulley transmission ratio of 1.5. The clutch system highlighted in Fig. 9, connected to the stationary side as mentioned in section II, to restrict the spindle system of the electromechanical actuator from rotating. As shown in Fig. 12 (a), a load cell (Kistler strain gauge meter4703) is attached at the end of the drawbar, to measure the thrust produced by the drawbar by transforming the rotary motion of the drive motor, transmitted through the pulleybelt mechanism into a linear motion. A Yaskawa inverter (A1000) is used to perform the speed control of the drive motor",
            " Unlike the conventional hydraulic lathe, for the proposed electromechanical actuator, the clamping system does need simultaneous supply of power when changed to spindle mode. The self locking produced by the screw-thread structure of the drawbar can maintain the clamping force even though the power is cut and transfer to the spindle mode. This ensures energy saving of the whole system by preventing simultaneous energy consumption by both modes. To perform the spindle operation using the experimental setup shown in Fig. 14 (a), the drive motor is operated in high speed mode and the clutch system shown in Fig. 9, is attached with the motor side plate as mentioned in section II. The speed plot showing the speed variation between drive motor and the spindle is recorded using two ONOSOKKI-FT-7200 techometers installed as shown in Fig. 14(a). The recorded results are presented in Fig. 14(b). The results in Fig 14(b) ensures the capability of the electromechanical actuator to produce a spindle speed of 4000 rpm when the drive motor speed is 6,000 rpm at a pulley transmission ratio of 1.5. A detailed design of a novel single drive motor integrated electromechanical actuator for the CNC lathe machine operation is presented in this paper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000318_icrae48301.2019.9043822-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000318_icrae48301.2019.9043822-Figure2-1.png",
        "caption": "Figure 2. The simulation model of six-legged robot in V-rep and leg mechanism",
        "texts": [],
        "surrounding_texts": [
            "There are many traditional methods based on physical model[7,8,9,10]. The most popular one is the modular controller design, which is to divide the robot leg control into multiple sub-control modules and solve the problem by multistep problem solving. In the model-based approach, in order to achieve optimal results, the design of the next controller module is often performed after the modules of each subcontroller achieve local optimal results. For example, the robot is first approximated to a mass point, and then based on this, a module is used to calculate the position of the robot\u2019s landing point at the next moment. After the landing point is given, the next module calculates the moving parameter trajectory for the leg to be moved according to the landing point. The last module can use the PID controller to track the trajectory. In this method, the optimal solution for each module can be obtained first, and then the parameters are passed to the next module. B. Orbit Optimization Method In order to solve the problems encountered by the physicalbased method, a new orbit optimization method is proposed[11]. In the orbit optimization method, the robot first plans the best path to the target, and then controls the robot to reach the target along the optimal path. In the orbit optimization method, the layered control method is often adopted, the top controller is designed for planning, and the bottom controller is used for tracking, and the two cooperate to control the robot to reach the target position. C. Deep Reinforcement Learning Reinforcement learning provides robots with a tool set and framework for designing complex and artificially difficult engineering behaviors. Its application in the field of robotics has been a common research topic in academia and industry. Previously, researchers at the University of Edinburgh developed a layered framework based on DRL to obtain a variety of human form balance control strategies. Their framework was pre-released on arXiv and presented at the 2017 International Conference on Humanoid Robotics[12]. The paper says that robots can achieve better balancing behavior than traditional controllers. Their research focuses on the use of DRL to solve the dynamic motion of humanoid robots. In the past, motion was mainly done using traditional analytical methods, that is, based on models, but now if you want to join the human form mechanism, then the robot needs higher computational processing power while DRL is appropriate. Another advantage of using DRL is that the robot\u2019s calculations can be processed offline, making the humanoid robot\u2019s online performance faster. In view of the increasingly powerful DRL algorithms, more and more research has begun to use DRL to solve control problems and get more diverse control strategies by using it."
        ]
    },
    {
        "image_filename": "designv11_14_0002856_s13369-021-05654-z-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002856_s13369-021-05654-z-Figure4-1.png",
        "caption": "Fig. 4 a CAD representation of printing orientation of DMLS plates; b represents the extraction of DMLS plates after printing",
        "texts": [
            " The plate was built with the three-layer system, i.e., upskin, core and downskin, and is shown schematically in Fig.\u00a03a. The two layers first printed at the bottom are called \u2018downskin\u2019; the three layers printed in the end are called \u2018upskin\u2019 and the rest of the layers between upskin and downskin are called \u2018core\u2019 of the AM 1 3 plate. Each layer following another in the regular sequence of the core was rotated by 67\u00b0 during printing, as shown in Fig.\u00a03c. The CAD representation of the printing orientation of DMLS plates is shown in Fig.\u00a04a, followed by the actual printed plates on the substrate base plate as shown in Fig.\u00a04b. Eight plates of additively manufactured aluminum alloy (AlSi10Mg) of size (100 \u00d7 50 \u00d7 3) mm were selected for post-processing using shot blasting. Each plate was further sectioned into four specimens of size (50 \u00d7 25 \u00d7 3) mm (Fig.\u00a01a) using WEDM (Wire-cut electric discharge machining). The sectioned specimen plates were cleaned in acetone to remove any debris or loose particles present on the surfaces before subjecting them to post-processing. Shot blasting of the test specimens was carried out using spherical glass beads with different particle sizes ranging between 0 and 150\u00a0\u03bcm, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001019_rpj-07-2018-0182-Figure18-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001019_rpj-07-2018-0182-Figure18-1.png",
        "caption": "Figure 18 Fabricated left and right die after support removal",
        "texts": [
            " Increase in fabrication time can be attributed to the following reason, the outer shells (red lines) that represent the exterior surface of the object were printed at a slow speed compared to infill (green lines) to maintain the dimensional accuracy of the part being Figure 13 Static analysis \u2013 stress distribution Figure 15 Stress distribution in the redesigned die Metal bellow hydroforming Prithvirajan R. et al. Rapid Prototyping Journal D ow nl oa de d by N ot tin gh am T re nt U ni ve rs ity A t 0 2: 47 3 1 M ay 2 01 9 (P T ) printed. The unidirectional pore shapes add a number of outer shell contours in every layer which accounts for increase in fabrication time. During post-processing, support structures were removed from the die. Figure 18 shows the fabricated die. Metal bellow is formed from 0.3 mm thick SS304 tube with redesigned ABS die in 25Ton hydroformingmachine at 9MPa. The first convolution is formed at 6 MPa hydraulic pressure, which is used to locate tube for forming next convolution. Then pressure is raised to 9MPa and next 10 convolutions are formed consecutively. Formed Bellow is shown in Figure 19. It is also observed that width of the convolution varies with the die opening height during the bulging stage. Further, the same die was used to form 20more convolutions without die failure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001733_lra.2020.3003235-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001733_lra.2020.3003235-Figure4-1.png",
        "caption": "Fig. 4. Angular Momentum About Impact Foot showing Pre- and Post-Impact Configurations. (a) Pre-Impact State x\u2212. (b) Post-Impact State x+.",
        "texts": [
            " The angular momentum L for a system can be given by L =mrcom \u00d7 (\u03c9 \u00d7 ro) (6) Where rcom = [ \u2212l sin(\u03b8) l cos(\u03b8) 0 ] is the vector from point of rotation to the COM, ro is the vector from the touchdown point to the COM, and \u03c9 = [ 0 0 \u03b8\u0307 ]T is the rotation rate vector. Applying (6) to the system shown in Authorized licensed use limited to: University College London. Downloaded on July 04,2020 at 10:27:22 UTC from IEEE Xplore. Restrictions apply. 2377-3766 (c) 2020 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. Fig. 4, the pre- and post-collision angular momentum can be given by L\u2212 =m \u2212l sin \u03b8l cos \u03b8 0 \u2212 \u00d7  0 0 \u03b8\u0307 \u2212 \u00d7 l sin(\u03b8 + \u03c6) l cos(\u03b8 + \u03c6) 0 \u2212  = [ 0 0 ml2 cos(\u03c6\u2212)\u03b8\u0307\u2212 ]T (7) L+ =m \u2212l sin \u03b8l cos \u03b8 0 + \u00d7  0 0 \u03b8\u0307 + \u00d7 \u2212l sin \u03b8l cos \u03b8 0 +  = [ 0 0 ml2\u03b8\u0307+ ]T (8) By equating (7) and (8), we can relate the velocity before impact \u03b8\u0307\u2212 to the velocity after impact \u03b8\u0307+. Additionally, through the kinematics seen in Fig. 4, we can relate the pre- and post-collision configurations, (\u03b8+ = \u03b8\u2212 + \u03c6\u2212) and (\u03c6+ = \u2212\u03c6\u2212). Together, these kinematic and momentum relationships form the discrete impact map function M x+ = M(x\u2212)\u03b8\u03c6 \u03b8\u0307 + = 1 1 0 0 \u22121 0 0 0 cos\u03c6\u2212 \u03b8\u03c6 \u03b8\u0307 \u2212 . (9) It is important to note that since the swing leg is considered massless and can be placed arbitrarily, the inter-leg angle \u03c6 does not appear in the continuous dynamics (5) but does appear in the discrete impact map (9). Based on the dynamics of a LIP, capture point (CP) [12] was developed as a means of controlling a force disturbance",
            " Instead of just a theoretical means of stopping motion, the CP can be defined as a boundary between walking and falling backwards as shown in Fig. 5. Computing the NIP CP is relatively straight forward, but there is no analytical solution. It begins with an understanding that capture occurs when the post-collision energy is equal to the max PE of the system (E+ = Umax) when the NIP is stationary and vertical (see Fig. 5). The local change between foot origin horizontal position \u2206x and height \u2206h in the global frame (xo, ho) (see Fig. 4) is described by \u2206x = x+ o \u2212 x\u2212o = l sin (\u03b8\u2212 + \u03c6\u2212)\u2212 l sin \u03b8\u2212 \u2206h = h+ o \u2212 h\u2212o = l cos \u03b8\u2212 \u2212 l cos (\u03b8\u2212 + \u03c6\u2212) (10) This section assumes that E refers to the local energy of a step (Sect. III-C introduces the concept of energy relative to some other reference point). Assuming we know the preimpact energy of a step E\u2212, we need to determine foot placement to dissipate all of the KE at the unstable equilibrium point T = \u03b8\u0307+ = 0 and E+ = U . Using (5), we can determine Authorized licensed use limited to: University College London"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001752_j.mechmachtheory.2020.103992-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001752_j.mechmachtheory.2020.103992-Figure13-1.png",
        "caption": "Fig. 13. Applied loads of the finite element model.",
        "texts": [
            " Stress analysis Based on finite element analysis, a loaded contact analysis can be completed to evaluate the stress distributions of the pinion and gear in Fig. 11 . The finite element model is as shown in Fig. 12 , which is meshed by the first-order hexahedral element. For the elements on the tooth surface, the max size of element length is set as 0.2 mm, and for other elements, the max size is set as 2 mm. The Young\u2019s modulus is 2.1 \u00d7 10 11 Pa and Poisson\u2019s ratio is 0.267. The torque of 540 Nm applied on the gear and the pinion is fixed as shown in Fig. 13 . The results of finite element analysis are shown in Fig. 14 . It can be seen that the shape and orientation of the contact ellipse of the finite element analysis are similar to those of the analytical analysis in Section 5.3 , which were used to verify the analytical model. Further, Fig. 14 also shows that the contact stress of the convex side of pinion is much larger than that of the concave side. This result has been predicted in Section 5.3 , because the length of the minor axis of the contact ellipse on the convex tooth surface of the pinion is generally less than that on the concave tooth surface of the pinion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001689_j.optlastec.2020.106325-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001689_j.optlastec.2020.106325-Figure2-1.png",
        "caption": "Fig. 2. Sample preparation: (a) Sample schematic diagram; (b) Laser processing system.",
        "texts": [
            " The 20CrMnTi was consisted of ferrite (F) and pearlite (P), with an average microhardness of 293.1HV0.3. The material steel of the ball joint cage of the steering axle was cut into pretreated sample with the size of 120 mm \u00d7 15 mm \u00d7 6 mm by DK7732 numerical control WEDM machine which was made by East China LTD.CO. The sandpaper was used to remove the cutting marks and oil traces on the surface of the sample to make it smooth and clean. A mesh unit with a pitch of 5 mm was processed on the sample by laser surface strengthening technique in Fig. 2(a). The unit specimen was treated in a laser processing system with a six-degree-of-freedom manipulator, as shown in Fig. 2(b). The system was equipped with Nd: YAG laser, cooling system, two-dimensional turntable, six-degree-of-freedom manipulator, servo control system and protective cylinder. Before laser processing, the movement of manipulator arm in Z axis direction of the worktable is controlled by servo control system. The amount of defocus of the laser lens was controlled to change the spot size. During the machining process, the servo control system was programmed to control the manipulator arm to drive the lens in the direction of the X-axis and the Y-axis of the worktable, so that units with different shapes, sizes, spacings and distribution directions could be obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002781_j.matpr.2021.02.045-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002781_j.matpr.2021.02.045-Figure2-1.png",
        "caption": "Fig. 2. Fixed Support.",
        "texts": [
            " In the form of a rectangle, the ends of the car bumper are formed. Table 1 Material properties of glass fiber and hybridized glass natural fiber reinforced composites. S.No Material Young\u2019s modulus GPa Poisson ratio Density Kg/m3 1 Glass fiber epoxy hybrid composites 45 0.25 2100 2 Hybridized glass fiber untreated Hemp fiber reinforced epoxy composites 62 0.28 1800 3 Hybridized glass fiber treated Hemp fiber reinforced epoxy composites 58 0.26 1600 The load is loaded into the front of the bumper. The Fig. 2 and Fig. 3 represents the boundary condition of the bumper. The meshing carried out here on the four wheeler bumper is a triangular mesh and the level of meshing is a fine mesh for the results will be as accurate as possible. The Fig. 4. represent the meshing of the bumper. Table 2 presents the Meshing details of glass fiber and hybridized glass natural fiber reinforced composite model bumper. Finite Element Analysis is the practical use of the Finite Element Method (FEM), which designers and researchers use to describe and solve complex primary, liquid, and multi-material science problem situations numerically and mathematically"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001366_tii.2019.2958818-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001366_tii.2019.2958818-Figure15-1.png",
        "caption": "Fig. 15. Rescue of two delivery agents; Coordinates are in meters.",
        "texts": [
            " Adversary begins its pursuit as its vision sensor identifies the delivery agent at D\u2032. In this scenario, the delivery agent is successfully rescued at T before a capture by the adversary. In a scenario where the vision of the adversary is obstructed by one of the agents, the adversary adopts an evasive maneuver to restore its vision. Fig. 14 illustrates this experiment. Here, the higher speed of adversary allows it to capture the delivery agent although it was initially obstructed by a rescue agent. Fig. 15 illustrates the pursuit of adversary amidst multiple agents. Ultrasonic sensors are employed to determine the closest delivery agent. The delivery agents are detected by the adversary (and vice-versa) as they arrive at D\u2032 1 and D\u2032 2. Rescue agent R2 employs Algorithm 6 to identify the weaker delivery agent as D1 and assigns itself for rescue. It, then, computes (and communicates) the meeting location T2 for R1 and D2. The adversary pursues until both delivery agents are rescued. Fig. 16 presents a scenario where, at the point of detection, the adversary is between two delivery agents (located at D\u2032 1 and D\u2032 2)",
            " It has been shown that the proposed algorithms allow quick computation of the meeting location, thus assisting in rescue of the delivery agent. The meeting point does not necessarily lie on the line joining locations of the delivery and rescue agents. Furthermore, the allocation of rescue agent to a given delivery agent is not based on its proximity to the latter but is dependent on its proximity to the meeting location computed via Algorithm 6. This follows from the fact that relative distances of rescue agents from adversary are captured via the safe regions. Consequently, in Fig. 15, the weaker delivery agentD1 is assigned toR2 and notR1, although R1 is closer to D\u2032 1 than R2. In this article, we studied the problem of safeguarding a vehicle carrying food/medicines from predatory attacks. We developed algorithms to determine capture/rescue of the delivery agent based on the notion of Apollonius circle. We showed that an autonomous adversary can successfully capture the delivery agent based on only a vision sensor and without communication hardware. Furthermore, explicit vision support on the delivery or rescue agents is not required for successful rescue"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002421_tmag.2020.3014789-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002421_tmag.2020.3014789-Figure2-1.png",
        "caption": "Fig. 2. Structure of WFSM studied in this paper. End plate functions as part of paths for armature reaction magnetic flux.",
        "texts": [
            " Downloaded on September 08,2020 at 08:54:46 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2020 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. HH-06 2 A. Wound Field Synchronous Machine Studied in This Paper In this paper, a 6-pole WFSM with rating of 145 V-20 Hz-2 kW was studied. The WFSM consists of a stator with a threephase armature winding and a salient pole rotor with a field winding as shown in Fig. 2(a) and (b). The field current is supplied to the rotor through a slip ring and brush. Both the stator core and the rotor core are composed of stacked steel sheets. The stacking length of the stator core and that of the rotor core are the same, 125 mm, but the rotor has end plates, which are 5 mm-thick solid steel, at the axial ends as shown in Fig. 2(b). Thus, the magnetic flux can flow not only in the cores but also in the end plates. This means that the effective stacking length of the rotor is longer than that of the stator. As shown in Fig. 2(b), the d-axis is defined as the direction of the magnetic saliency parallel to the direction of the magnetomotive force of the field winding. The q-axis is defined as the direction that is electrically perpendicular to the d-axis. B. Boundary Condition for Analysis of End Inductance Fig. 3(a) schematically shows a side view of the WFSM. In this paper, the end windings of the armature and field and the surrounding space are defined as an end region, and the analysis simulating only the end region is called a partial 3D FEA"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000205_srin.201900449-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000205_srin.201900449-Figure1-1.png",
        "caption": "Figure 1. Fracture toughness samples produced along the three different building direction: P|| where the BD is parallel to the notch; P\u22a5 where the BD is perpendicular to the notch; L where the BD is longitudinal to the notch.",
        "texts": [
            " The fracture toughness samples were produced along three different building directions. The specimen, according to ASTM E399, had a length of 30mm, a width of 3mm, and a thickness of 6mm. It was extracted by electron discharge machining (EDM) from a bigger specimen having dimensions of 30 30 6.5mm3. On the sliced samples, a notch of 3mm depth and 50 \u03bcm curvature radius (\u03c1) was produced by wire EDM. The notch was cut parallel (P||), perpendicular (P\u22a5), and longitudinal (L) to the building direction (Figure 1). The samples were subjected to two different vacuum heat treatments, namely, QT and tempering only (T). Quenching was conducted at 1020 C for 15min followed by 5 bar Nitrogen cooling. Double tempering for 2 h\u00fe 2 h was performed at 625 and 650 C for quenched and AB samples, respectively. Selecting different temperatures was meant to achieve nearly similar hardness for both QT and T samples according to tempering curves.[6] Microstructural analysis was conducted by light optical and scanning electron microscopy (SEM)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003900_jsvi.1997.1323-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003900_jsvi.1997.1323-Figure1-1.png",
        "caption": "Figure 1. Schematic of the active journal bearing.",
        "texts": [
            " These active devices include: magnetic bearings [4], piezoelectric bearing pushers [5, 6], hydraulic actuator journal bearings [7], variable impedance bearings [8\u201310], damper using electro-rheological fluids [11], deformable bushes [12], and active journal bearing with a flexible sleeve [13]. This paper presents a modelling technique of multi-bearing rotor system incorporating the newly developed active journal bearing presented by Krodkiewski and Sun [13]. Both a general non-linear model and a linearization method are presented with numerical solutions and simulations. The flexible sleeve can be considered as a new feature of the proposed active journal bearing as shown in Figure 1. The sleeve is activated by the chamber pressure pc , which is controlled by valves in the hydraulic system. The oil film of the bearing and the pressure chamber is separated by the flexible seal. Therefore, the chamber pressure will not influence the boundary conditions of the oil film. The deformation of the flexible sleeve can be changed by adjusting the chamber static pressure. Therefore the geometry and thickness of the oil film, and hence the dynamic properties of the rotor system, can be controlled without stopping the operation of the machine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002578_j.jmatprotec.2020.117037-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002578_j.jmatprotec.2020.117037-Figure4-1.png",
        "caption": "Fig. 4. Forming process and result: (a) schematic diagram of laser cladding & laminated deposition cladding process; (b) morphology of the sample.",
        "texts": [
            " The inert chamber also included a forming system, a fiber-winding device, and a wire-laying device. The design route for the LCLD device that was developed by our laboratory is shown in Fig. 3. First, the SiC fiber was stored in the fiber-winding device. When manufacturing the composite, fiber was pulled from the fiber-winding device by the wire-laying device to form a fiber layer. Then, the SiC fiber layer was laid on the forming substrate and tightened. Finally, titanium alloy powder was deposited on the SiC fiber layer. The forming process is shown in Fig. 4a. The size of the forming composite was 30 mm x 20 mm on a titanium alloy forming substrate. The macroscopic appearance of the single-layer sample is shown in Fig. 4b. First, the effects of the defocusing amount, shield gas flow rate, laser power, powder feed rate, and scanning speed on the cladding quality were evaluated. Since studies on composite manufacturing using LCLD are scarce, there are no literatures that provide a reference for designing such an experiment. Therefore, a single-channel single-factor experiment was first conducted to determine whether these five process parameters affect the cladding quality to determine the approximate process parameter range (Table 3)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure30.3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure30.3-1.png",
        "caption": "Fig. 30.3 Husky robot URDF model and its transformations tree (left\u2014robot model visualization, right\u2014robot model links hierarchy)",
        "texts": [
            " On the contrary, uneven environment\u2019s floor represents a surface with changes in height of up to 1.3 m relative to the ground. Illumination is even all over the environment without shadows. In this work, we use Husky mobile robot provided by the husky_description ROS package and modify it to include necessary sensors. The model of the robot is represented in the unified robot description format (URDF). The model consists of links connected by joints with their visual and collision meshes. The simplified transformation tree and robot model are presented in Fig. 30.3. The robot is equipped with the following sensors: (a) stereo view camera (resolution: 640 \u00d7 480, focal length: 320 pixels, RGB); (b) laser range finder (field of view 270\u25e6); (c) inertial measurement unit (IMU). Gazebo sensors API enables to incorporate distortions into the sensory input. For example, attached cameras can have pixel intensity noise (e.g., Gaussian) as well as radial and tangential lens distortions. For stereo cameras, we employed a mutli-camera sensor API to provide synchronized global shutter images from both cameras [19]",
            " The average linear velocity is around 0.15m/s and angular\u20140.15 rad/s. Final trajectories do not have sharp turns to avoid drastic changes in the observed environment. In our datasets collection, we employ a system with the following properties: (a) Intel Core i7-8750H@2.2 GHz CPU; (b) 16 GB LDDR4 RAM; (c) 1TBM.2 PCI-E SSD; (d) Ubuntu 16.04 LTS OS (with ROS Kinetic installed); The ground-truth 6D pose trajectory is obtained using the P3D Gazebo plugin. Poses are provided for the base_footprint frame w.r.t the world fixed frame (Fig. 30.3). Datasets were recorded into ROS bag files, they contain the following topics: (a) /clock\u2014simulation timestamp; (b) /cmd_vel\u2014control commands; (c) /odometry/filtered\u2014IMU and wheel odometry EKF fusion; (d) /ground_truth\u20146D pose of the robot w.r.t world; (e) /tf_static\u2014transformations between robot links; (f) /scan\u2014LiDAR scans data; (g) /camera/left/image_raw and /camera/left/camera_info\u2014left stereo camera top- ics with image data and meta information; (h) /camera/right/image_raw and /camera/right/camera_info\u2014right stereo camera topics with image data and meta information; We propose to record datasets and distribute them with the simulated environment (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002590_cvci51460.2020.9338578-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002590_cvci51460.2020.9338578-Figure1-1.png",
        "caption": "Figure 1. Vehicle kinematic model",
        "texts": [
            "00 \u00a9 2020 IEEE 627 20 20 4 th C AA In te rn at io na l C on fe re nc e on V eh ic ul ar C on tr ol a nd In te lli ge nc e (C VC I) | 97 8- 1- 72 81 -8 49 7- 5/ 20 /$ 31 .0 0 \u00a9 20 20 IE EE | D O I: 10 .1 10 9/ CV CI 51 46 0. 20 20 .9 33 85 78 Authorized licensed use limited to: UNIVERSITY OF BATH. Downloaded on May 14,2021 at 08:10:23 UTC from IEEE Xplore. Restrictions apply. II. VEHICLE MODEL In this section, a kinematic vehicle model is employed to design the controller. The kinematic model of a two-wheel differential driving AGV is shown in Fig. 1. Assuming that the AGV is a rigid body, the longitudinal and lateral positions and the yaw angle of the vehicle can be described as follows: . . . cos( ), sin( ), c c c c c x v y v             (1) where v is the linear velocity, and  is the angular velocity. The width wR and length lR are the width and length of the AGV, respectively. ( cx , cy ) is the center point of AGV, and c is the angle between the AGV direction and the horizontal direction. Therefore, z = [ , , ]T c c cx y  forms the states of the robot",
            " If [ ]Tu v  is the control input of AGV, and lv and rv are the velocities of the left and right wheels, then lv and rv are bounded by max max | | , | | l r v v v v    (2) where maxv is the maximum linear velocity. The control inputs v and  can be expressed as functions of lv and rv as 2 2 l r r l v v v v v        . (3) According to Eqs. (2) and (3), the control input is limited as max| | | |v l v   , (4) where \ud835\udc59\u03c9 is half of the track width of two wheels. As presented in the motion equation in Fig. 1, the angular velocity r and l of the left and right wheels can be obtained according to v and \u03c9 and given as: / 2 / 2 / 2 / 2 l r r rv r l r l                    , (5) where r is the radius of the driving wheel. According to Eqs. (1) and (5), the kinematic model of the two-wheel differential AGV trolley can be described as . . . cos( )( ) / 2, sin( )( ) / 2, ( ) / 2 c c r l c c r l c l r x r y r r l                     (6) According to Eq. (6), the model of two-wheel differential AGV trolley can be built in Simulink"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000031_cdc.2015.7402793-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000031_cdc.2015.7402793-Figure1-1.png",
        "caption": "Fig. 1. Inertial landmarks on O as observed from vehicle S with optical measurements.",
        "texts": [
            " Additionally, this estimation scheme can be extended to relative pose estimation between vehicles from optical measurements, without direct communications or measurements of relative velocities. Consider a vehicle in spatial (rotational and translational) motion. Onboard estimation of the pose of the vehicle involves assigning a coordinate frame fixed to the vehicle body, and another coordinate frame fixed in the environment which takes the role of the inertial frame. Let O denote the observed environment and S denote the vehicle. Let S denote a coordinate frame fixed to S and O be a coordinate frame fixed to O, as shown in Fig. 1. Let R \u2208 SO(3) denote the rotation matrix from frame S to frame O and b denote the position of origin of S expressed in frame O. The pose (transformation) from body fixed frame S to inertial frame O is then given by g = [ R b 0 1 ] \u2208 SE(3). (1) Consider vectors known in inertial frame O measured by inertial sensors in the vehicle-fixed frame S; let \u03b2 be the number of such vectors. In addition, consider position vectors of a few stationary points in the inertial frame O measured by optical (vision or lidar) sensors in the vehicle-fixed frame S",
            " (37) For optical measurements, eight beacons are located at the eight vertices of the cube, labeled 1 to 8. The positions of these beacons are known in the inertial frame and their index (label) and relative positions are measured by optical sensors onboard the vehicle whenever the beacons come into the field of view of the sensors. Three identical cameras (optical sensors) and inertial sensors are assumed to be installed on the vehicle. The cameras are fixed to known positions on the vehicle, on a hypothetical horizontal plane passing through the vehicle, 120\u25e6 apart from each other, as shown in Fig. 1. All the camera readings contain random zero mean signals whose probability distributions are normalized bump functions with width of 0.001m. The following are selected for the positive definite estimator gain matrices: J = diag ( [0.9 0.6 0.3] ) , M = diag ( [0.0608 0.0486 0.0365] ) , (38) Dr = diag ( [2.7 2.2 1.5] ) ,Dt = diag ( [0.1 0.12 0.14] ) . \u03a6(\u00b7) could be any C2 function with the properties described in Section 3, but is selected to be \u03a6(x) = x here. The initial state estimates have the following values: g\u03020 = I, \u2126\u03020 = [0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001208_j.triboint.2019.105999-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001208_j.triboint.2019.105999-Figure5-1.png",
        "caption": "Fig. 5. Stress analysis diagram of the sealing ring.",
        "texts": [
            " The mechanical oblique-cone-slid-ring (OCSR) compensated seal designed by Hu et al. [6] The device can independently adjust the position of the axial bush, push the OCSR to squeeze the O-ring to deform and obtain greater contact stress, as shown in Fig. 4. However, in the actual working process of the piston rod seal, the seal pair is simultaneously subjected to the normal contact force FN for ensuring the tightness performance and the tangential frictional force Ff1 for resisting the reciprocating motion, as shown in Fig. 5. The increase of the normal contact force will cause the increase of the tangential friction force, which will aggravate the heating of the hydraulic cylinder and the wear of the seal pair. Therefore, it is necessary to find an inherent law between the compensating pressure and the friction characteristics of piston rod seals with different degrees of wear so as to better meet the sealing requirements without affecting the working efficiency of the hydraulic cylinder. However, this is still a problem that needs to be explored, so this paper has carried out related research on this issue"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002264_ab77d7-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002264_ab77d7-Figure1-1.png",
        "caption": "Figure 1.CADmodels of (a)Truncheon artefact and (b)Tyl\u00f6Helo sauna corner knot.",
        "texts": [
            " Most importantly, in this paper, a methodology is presented to identify the significant roughness parameters which can prove useful for quick assessment of the part quality. Finally, a case study is presented where the best finishing method and settings are implemented on a Tyl\u00f6Helo Company product and assessed for quality using advanced surface topography characterization technique. 2.Materials andmethods A truncheon artefact [10, 24]with varying build inclination from0\u02da to 90\u02da in steps of 10\u02da incrementwas designed for measuring the roughness at various slopes, as shown in figure 1(a). The sample was fabricated in a Stratasys uPrint by Dimension FDM printer with the ABS-P400 material, and the layer thickness was maintained at 0.254mm with low infill density setting. The print settings were held constant for all the artefacts produced. Figure 1(b) represents a CAD model of the component termed as corner knot produced by Tyl\u00f6Helo company. It is used as a stationary joint to hold together the aluminium frames that form the structure of a sauna room.This product is used as a case study in this paper. 2.2.Design of experiments The truncheon artefact facilitates in providing the roughness information at various angles, whichmakes it easy to visualize the roughness of a real industrial product without having to produce it. It also helps in saving the printing time and materials"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001453_1077546319900115-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001453_1077546319900115-Figure2-1.png",
        "caption": "Figure 2. Schematic diagram and sensor layout.",
        "texts": [
            " Step 5. The learned features from the training data are used to train the softmax classifier, and the testing data are used to validate the performance of the proposed method. Vibration data are collected from the RC of petroleum industry to validate the proposed method. A full description of the specific petroleum production company and a photo of the RC in operation could not be provided in this article because of the commercial-in-confidential nature, but a schematic diagram of the RC is shown in Figure 2. Figure 2 illustrates the structure of RC and the layout of sensors, where a RC with four cylinders is taken as an example. To provide more details about the instrument, Figure 2(a) and (b) are used to display the general structure and the inner structure from A-A cross section, respectively. The engine lying on the left side provides the output power to drive the rotation of the shaft. Phase sensors are installed on the flywheel near the engine to monitor the rotating speed. Accelerometers are attached on the crosshead to measure the vibrations of the cylinder. The signals collected by the accelerometer can reflect the operating conditions of the RC. The piston rods in the cylinder can take in and push out gas via the gas valve"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002021_s11661-020-06076-6-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002021_s11661-020-06076-6-Figure1-1.png",
        "caption": "Fig. 1\u2014Schematic diagram of LPBF process: (a) SLPBF; (b) MLFA-PBF; (c) MLA-PBF (the laser beam arrangement direction is perpendicular to the scanning direction); (d) MLA-PBF (the laser beam arrangement direction is parallel to the scanning direction).",
        "texts": [
            "[3,4] In view of this, multi-laser powder bed fusion (MLPBF) has gradually gained attention in recent years, and it has become one of the most promising technologies for rapid manufacturing of large metal parts.[5,6] MLPBF technology is mainly divided into two categories: multi-laser forming area powder bed fusion[7] (MLFA-PBF) and multi-laser array powder bed fusion[8] (MLA-PBF). Among them, MLFA-PBF technology divides the powder bed into multiple forming zones, and then uses multiple laser beams to act on their respective forming zones simultaneously. Compared with the SLPBF process (Figure 1(a)), MLFA-PBF is equivalent to reducing the area of the forming zone for a single-laser beam (Figure 1(b)). MLFA-PBF generally uses two or four laser beams and uses fiber lasers.[9] MLA-PBF technology puts multiple laser beams side by side, and then acts on the powder bed at the same scanning speed. Compared with the SLPBF process, MLA-PBF is equivalent to widening the width of the laser action area (Figure 1(c)), or increasing the number of laser beams acting on a single solidified track (Figure 1(d)). MLA-PBF does not limit the number of laser beams (the number of laser beams in Reference [10] has reached sixteen), and the laser can be a diode laser[10] or a fiber laser.[11] The current MLPBF experimental researches mainly focus on the surface Liu Cao is with the School of Mechanical and Electrical Engineering, Guangzhou University, Guangzhou, 510006 China Contact e-mail: caoliu@gzhu.edu.cn. Manuscript submitted August 31, 2020; accepted October 18, 2020. METALLURGICAL AND MATERIALS TRANSACTIONS A roughness, porosity and microhardness of the parts",
            " Especially for the MLA-PBF process, the molten pool dynamics when multiple laser beams act side by side on the powder bed has an important influence on the optimization of the MLA-PBF process, which is also the focus of this paper. In this paper, the MLA-PBF spreading powder process was calculated based on the open source DEM framework Yade, and the dynamic behavior of the MLA-PBF molten pool was described based on the open source CFD framework OpenFOAM. For the single-line mode of MLA-PBF (multiple laser beams forming the same solidified track, Figure 1(d)) and the multi-line mode of MLA-PBF (multiple laser beams simultaneously forming multiple solidified tracks, Figure 1(c), the influences of laser power and laser beam space on the forming process were simulated and METALLURGICAL AND MATERIALS TRANSACTIONS A verified by comparison with experimental results. This paper is expected to provide theoretical support for deepening the application of MLA-PBF in metal additive manufacturing. The premise of describing the molten pool dynamics in the MLA-PBF process based on the mesoscopic scale is to obtain the particle distribution of the powder bed. In the spreading powder process, the metal particles are pushed and squeezed by the roller, so a corresponding particle dynamic model needs to be established to describe the mechanical action between the roller and the particles, as well as the particles and the particles",
            " Figure 5 shows the cumulative probability curve of the particle size herein. METALLURGICAL AND MATERIALS TRANSACTIONS A The physical property parameters of 316L stainless steel required to predict the MLA-PBF process based on OpenFOAM are shown in Table I. The computing resource configuration used was Intel Xeon Gold 5120 CPU (dual CPU, 56 threads, and 96 GB memory). The single-line mode was first focused, that is, considering the situation where multiple laser beams form the same solidified track (Figure 1(d)). Figure 6 shows the particle distribution of the powder bed in the Table I. Physical Property Parameters of 316L Stainless Steel Parameter Value Unit Density of Metal 7270 kg/m3 Specific Heat of Metal 790 J/(kg\u00c6K) Thermal Conductivity of Metal 24.55 W/(m\u00c6K) Solidus Temperature 1658 K Liquidus Temperature 1723 K Evaporation Temperature 3090 K Latent Heat of Melting 2.7 9 105 J/kg Latent Heat of Gasification 7.45 9 106 J/kg Viscosity of Liquid Metal 0.00345 Pa\u00c6s Surface Tension 1.6 N/m Temperature of Surface Tension 8 9 10 4 N/(m\u00c6K) Molecular Mass 9",
            " When the laser beam space was too large, the metal particles in a certain position were preheated and pre-sintered under the action of the front-laser beam; however, because the time required for the rear-laser beam to reach this position was too long, the metal-phase temperature in front of the molten pool was not high, causing the front-laser beam to lose its preheating function. It can be seen that a moderate laser METALLURGICAL AND MATERIALS TRANSACTIONS A beam space should be used for the dual-laser forming of low-front and high-rear. The optimal laser beam space obtained herein is the value of the spot diameter. The multi-line mode was then focused, that is, considering the situation where multiple laser beams simultaneously form multiple solidified tracks (Figure 1(c)). Figure 13 shows the particle distribution of the powder bed in the multi-line mode, where the geometric dimensions of the substrate are as follows: X-direction (1000 lm), Y-direction (400 lm), and Z-direction (50 lm), and the thickness of the powder bed is 45 lm. The geometric dimensions of the mesh model are as follows: X-direction (1000 lm), Y-direction (400 lm), and Z-direction (145 lm), and the mesh size is 2.5 lm. Table III shows the calculation schemes in the multi-line mode, where the laser beams all moved from the X-coordinate of 50 lm to the X-coordinate of 950 METALLURGICAL AND MATERIALS TRANSACTIONS A lm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002471_j.engfailanal.2020.104811-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002471_j.engfailanal.2020.104811-Figure12-1.png",
        "caption": "Fig. 12. Excitation load applied to the CDBG, where ng, g are the teeth number and operational speed of the driven gear, r is the operational speed of the gas generator rotor and nr is the possible multiplier; Fh,max is the amplitude of the high-frequency excitation force; Fl0 and Fl are the mean value and amplitude of the low-frequency excitation force.",
        "texts": [
            " The excitation frequency of gear-teeth meshing Hh can be expressed as the product of the operational speed of the CDBG g and the number of the gear teeth ng: =H n \u00b7h g g (4) Because the frequency of gear-teeth meshing is quite high and the meshing time is short (less than 1 ms), it is necessary to consider the high-frequency excitation as an impact load. For some more practical reasons, treat these high-frequency excitations as impact load can also reduce time steps significantly in the transient dynamic analysis of the CDBG. Therefore, the high-frequency excitation load applied at the CDBG can be approximated as an impact function (Fig. 10(a)), whereas the low-frequency excitation load can be represented by a sine function (Fig. 12(b)). The amplitude of the lateral vibration of the gas generator rotor is significant due to its high operational speed and weak front bearing stiffness. The rotor vibration will therefore stimulate an excitation on the CDBG. Since the contact stiffness of the gear pair in an actual gear-rotor system has strong nonlinearity, the excitation frequency applied to the CDBG by the vibrating rotor will contain multiple frequencies, which can be expressed as =H n \u00b7 ,l v r (5) where r is the operational speed of the gas generator rotor and =n 1, 2, 3. ..v is the possible multiplier. The multiple frequencies generated by lateral vibration of the gas generator rotor are shown to exist experimentally by the vibration signal analysis illustrated in Fig. 5. This period of excitation is related to the operational speed of the gas generator rotor. This excitation has a lower frequency and a larger amplitude than the excitation produced by gear teeth meshing, and can be expressed by a harmonic function. This is shown in Fig. 12(b), where the initial force exerted at the interface is Fl i, , and the amplitude of the variable force is Fl. (2) Campbell diagram and vibration response of the CDBG with multi-source excitation Since the external excitation acting on the CDBG has two sources, the Campbell diagram, which is based on a single excitation system, is not applicable here. Fig. 13 shows a modified Campbell diagram of the CDBG, in which both the 43\u00d7 excitation, based on the operational speed of the CDBG, and the 2\u00d7 excitation, based on the operational speed of the rotor, are considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000202_012024-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000202_012024-Figure1-1.png",
        "caption": "Figure 1. Design Reasearch Part (a)Test-rig design. (b)KARLING\u2019s design on Autodesk Inventor",
        "texts": [],
        "surrounding_texts": [
            "The BLDC motor used is 1000W 205 40H BL1 BLDC Electric Bike Bicycle Spoke Motor Hub produced by QS MOTOR. 1000W 205 40H BL1 BLDC Electric Bike Bicycle Spoke Hub is used to be able to support car loads and could reach a maximum speed of 50 km/h. The controller used is the KBL48151X product from Kelly Controller. The controller can support 24 - 48V voltage sources with a maximum current of 150A and is used to control BLDC motors. The battery used has a voltage equal to the voltage to drive the motor and a capacity of 20Ah. The battery circuit is designed using batteries with lithium-ion type."
        ]
    },
    {
        "image_filename": "designv11_14_0001806_acc45564.2020.9147931-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001806_acc45564.2020.9147931-Figure2-1.png",
        "caption": "Fig. 2. Notation used for modelling the Convergence UAV.",
        "texts": [
            "00 \u00a92020 AACC 4175 Authorized licensed use limited to: Cornell University Library. Downloaded on August 19,2020 at 02:45:50 UTC from IEEE Xplore. Restrictions apply. the rigid body equations of motion are given by p\u0307 = Rv (2) v\u0307 = \u03c9\u00d7v + gR>e3 + 1 m F (3) R\u0307 = R\u03c9\u00d7 (4) \u03c9\u0307 = \u2212J\u22121\u03c9\u00d7J\u03c9 + J\u22121M, (5) where e3 = (0, 0, 1)> [6]. The Convergence UAV has seven actuators: two elevons, three rotors, and two servos which control the angle of the front rotors. The rear rotor is fixed and points in the \u2212zB direction of the aircraft body frame. As shown in Fig. 2, the elevon deflection values are denoted by \u03b4e = (\u03b4e1 , \u03b4e2)T where \u03b4e\u2217 \u2208 [\u2212\u03b4emax , \u03b4emax ] and \u03b4emax is the maximum elevon deflection in radians. Additionally, the throttle inputs are denoted by \u03b4r = (\u03b4r1 , \u03b4r2 , \u03b4r3)T where \u03b4r\u2217 \u2208 [0, 1], and the rotor servo angles are denoted by \u03b8r = (\u03b8r1 , \u03b8r2 , \u03b8r3)T . The range for \u03b8ri has been measured to be \u03b8r1 , \u03b8r2 \u2208 [0, 7\u03c0/12] radians, and \u03b8r3 = \u03c0/2 radians. We model the equations of motion for the rotation of the front two rotors using the first order dynamics \u0398\u0307 = ks(\u0398c \u2212\u0398), (6) where ks \u2208 R defines the rise time, \u0398 = ( \u03b8r1 , \u03b8r2 )> , and \u0398c = ( \u03b8cr1 , \u03b8cr2 )> are the angles commanded to the servo motors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002731_j.mechmachtheory.2021.104285-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002731_j.mechmachtheory.2021.104285-Figure7-1.png",
        "caption": "Fig. 7. GTS of different sizes.",
        "texts": [
            " \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 \u03bb\u03c9 = 1 \u03bbt = 1 \u03bb\u03b7 \u03bbX = \u03bbF \u03bbK = \u03bbF \u03bb\u03b7 \u03bb \u02d9 X = \u03bbF \u03bbC = \u03bbF \u03bb\u03b7 2 \u03bbX\u0308 = \u03bbF \u03bbM = \u03bbF \u03bb\u03b7 3 \u03bbF = \u03bbT \u03bb\u03b7 = \u03bbP \u03bbn \u03bb\u03b7 = \u03bbF m = \u03bbP (69) The similarity ratio of the dynamic response in the translation direction is written in Eq. (70) . \u23a7 \u23aa \u23a8 \u23aa \u23a9 \u03bbx = \u03bby = \u03bbz = \u03bbF \u03bbK = \u03bbF \u03bb\u03b7 \u03bb \u02d9 x= \u03bb \u02d9 y= \u03bb \u02d9 z= \u03bbF \u03bbC = \u03bbF \u03bb\u03b7 2 \u03bbx\u0308 = \u03bby\u0308 = \u03bbz\u0308 = \u03bbF \u03bbM = \u03bbF \u03bb\u03b7 3 (70) The similarity ratio of the dynamic response in the torsion direction is written in Eq. (71) . \u23a7 \u23aa \u23a8 \u23aa \u23a9 \u03bb\u03b8x = \u03bb\u03b8y = \u03bb\u03b8z = \u03bbF \u03bb\u03b7 2 \u03bb \u02d9 \u03b8x = \u03bb \u02d9 \u03b8y = \u03bb \u02d9 \u03b8z = \u03bbF \u03bb\u03b7 3 \u03bb = \u03bb = \u03bb = \u03bbF (71) \u03b8\u0308x \u03b8\u0308y \u03b8\u0308z \u03bb\u03b7 4 Combined with the finite element method, the parallel shaft GTS is modeled. The full size model and different scale models are shown in Fig. 7 . All three models can be transformed into the finite element model shown in Fig. 2 . The gear parameters, mesh stiffness, cumulative pitch error and tangential error of single tooth are listed in Table 1 . The dimension parameters and material parameters of shaft segments are listed in Table 2 . L j and D j ( j = 1~18) are the length and outer diameter of the shaft segment, respectively. After the structural parameters of the system are determined, the similarity ratio of the dynamic response of the system can be obtained based on the scaling law derived in Section 2 , as shown in Table 3 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003877_1.2833881-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003877_1.2833881-Figure8-1.png",
        "caption": "Fig. 8 Three-dimensional finite element model",
        "texts": [
            " Specifically, there were 200 horizontal divisions along each coating layer, making the elements within the coating 0.1905 mm wide. Utilizing the two-dimensional model as a basis, ball, plate, and coating volumes were generated by rotating the right half of the cylinder-plate system about a 180 degree arc. The re- Journal of Tribology OCTOBER 1997, Vol. 1 1 9 / 7 5 7 Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use suiting three-dimensional model is depicted in Fig. 8. As shown in Fig. 9, the three-dimensional coated ball system uses physical symmetry about the z-axis. Only symmetry along the balls roll ing path could be incorporated due to the tangential friction forces which develop in the model. The ball, plate and coatings all consisted of eight-noded, three-dimensional solid elements which could be defined with up to nine orthotropic material constants (Eu, \u00a3'22, \u00a333, G12, Gn, G23, 1\u0302 12. t^n, and 1/23)- At the completion of the three-dimensional solid element genera tion, the model depicted in Fig. 8 consisted of a total of 44,825 ball, plate, and coating elements. Once the ball, plate, and coating elements were created, the model was placed into contact. This was accomplished using a three dimensional point-to-surface contact element which were used to represent contact between two surfaces. Each contact element had five nodes with three degrees of freedom ( and Mj) at each node. As described in Section 2.1, contact occurs when a contact node penetrates a target surface. Therefore, in the three dimensional model, the finely meshed external coating nodes were chosen as the target surfaces, while the points along the top and bottom portions of the sphere were selected as the contact nodes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001968_j.icheatmasstransfer.2020.104868-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001968_j.icheatmasstransfer.2020.104868-Figure2-1.png",
        "caption": "Fig. 2. Blade casting (a), standard (b, c) and modified (d, e) mold with connected thermocouples, ring radiation baffle (SRB) (f), adjusted radiation baffle (ARB) (g), ARB and inner radiation baffle (IRB) (h), schematic view of the assembly of inner radiation baffles (i), 1\u20139 are test points of temperature measurement.",
        "texts": [
            " Therefore, the impact of radiation baffle geometry on thermal radiation shielding and microstructure refinement was investigated in the paper and compared with the results obtained for blades produced using a new IRBs technique. The directional solidification of single crystal blades were carried out using the industrial-scale Bridgman method by withdrawing the mold from the heating area through the opening in the radiation baffle to the cooling area of the furnace at rate 3 mm/min (Fig. 1). In order to conduct the experiments, ceramic molds and then single crystal castings were produced (Fig. 2a-d). The wax assembly contained 5 dummy blades 140 mm high and 60 mm wide, which consisted of upper and lower platforms, root and airfoil (Fig. 2a). Three wax assemblies and ceramic molds with the same design were made (Fig. 2b-d). The temperature measurement was carried out in points 1\u20139 using type B thermocouples located along the blade height at the distance of 22, 29, 37, 48, 65, 82, 99, 116 and 127 mm from the casting base (Fig. 2a, b). However, in next two experiments, the temperature measurement was carried out at points 3\u20137 as shown in Fig. 2a, c, d. The precise dimensions and shape of these single crystal blades as well as the preparation of the mold and temperature measurement are given in previous investigations [20-22]. Each of the molds was preheated to the temperature of 1520 \u00b0C and then the molten CMSX-4 Ni-based superalloy of the same temperature and in the amount of 3.5 kg was poured into the mold cavity. Three experiments, widely differing in the efficiency of baffle separating in the furnace, were performed. In the first experiment, the standard mold and the ring-shaped radiation baffle (SRB) with the opening diameter of 230 mm (Fig. 2f) were used. The baffle wasThe baffle was placed on the thermal insulation of the heating chamber (Fig. 1a). In the second experiment, the opening of radiation baffle was adjusted as much as possible to the largest cross-section of the blade and also to the pouring cup of the standard mold (Fig. 2 g). This baffle will be called adjusted radiation baffle (ARB) in the following part of this study. The same geometry of ARB and new modified design mold were used in the next process (Fig. 2d, h). The mold was modified by introducing the inner radiation baffles (IRBs) between the central rod and blades (Fig. 2d). In order to place the inner radiation baffles, part of the central rod was first removed from the standard mold (Fig. 2e). Then, the alumina ceramic tube and several ceramic washers were inserted in the place of the removed central rod. The inner radiation baffle was formed by inserting graphite elements between the two ceramic washers as shown in Fig. 2i. Each of the inner radiation baffle with diameter of 155 mm consisted of five graphite elements connected together and matched to the external shape of the mold (Fig. 2h, i). In that way, seven inner radiation baffles were mounted one by one at the distance of 60, 85, 108, 128, 148, 170 and 195 mm from the chill plate as shown in Fig. 2 i. One blade from each experiment was cut off at temperature measurement height (points 1\u20139) in order to analyze the dendritic microstructure on the cross-sections of casting. The temperature measurement was carried out at various points of the blade during preheating the mold, pouring liquid metal, solidification and cooling the casting (Fig. 3a-c). Typical cooling curves were obtained, which differed in shape, depending on the place of temperature measurement and the conducted process. Based on their analysis, exemplary temperature distributions along the blade height (profile 1\u20133) were obtained for the time, when liquidus isotherm is located at the height of test point 5 of the casting (Fig",
            "5 \u00b0C, when IRBs (profile 3) were used, compared with the temperature of the standard process with SRB, where it reached T6 = 1416.5 \u00b0C (profile 1). It was found that the change of temperature along the casting height increased when the contour of opening in the radiation baffle was the closest to the outer surface of mold. The smallest gap between the castings and central rod of mold, and therefore, the greatest thermal separation of heater from the cooling area were achieved for the process modified with IRBs (Fig. 2 h). In consequence, the thermal impact of heaters on mold surface, below the radiation baffle was limited and resulted in a reduction of the casting temperature to a greater extent than for the standard process with SRB or ARB. For the standard process, the opening contour of ARB was limited by the diameter of the pouring cup and the largest cross-section of the blade (Fig. 1b). The mold surface, being in the cooling area, was only partially separated from the heater area of the furnace, compared with the process, in which IRBs were additionally used",
            " On the basis of analysis of the obtained research results and those presented in references [19], the new design and a proposal of the implementation of IRBs technique for the production of large IGT blades using Bridgman method has been discussed in this article. Fig. 8b shows the shape and location of IRBs in the assembly, taking into account the same geometry of external radiation baffle as it used to be in the standard process (Fig. 8a). The geometry of IRBs and method of their location in the assembly differ from those shown in Fig. 2 for the production of small aero engine turbine blades. The IRBs were divided into sections (designated as IRBs 1\u20133) in order to match them more precisely to the shape of the blade (IRBs 1 and IRBs 2) and to position them easily in the assembly (IRBs 3) (Fig. 8b). The presented solution provides the reduction of the gap width by the IRBs, mainly at the airfoil height, simultaneously keeping the area of ARB as large as possible. However, the opening contour of external radiation baffle is sufficiently matched to the geometry of the starter and root"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002408_10667857.2020.1797284-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002408_10667857.2020.1797284-Figure2-1.png",
        "caption": "Figure 2. Diagram of WAAM process.",
        "texts": [
            "2 mm) provided by NEIMM and 6061 aluminium alloy plate with the size of 300 mm \u00d7 150 mm \u00d7 10 mm was selected as substrate. The shielding gas of WAAM process is argon gas with a purity of 99.999%, and the chemical composition of the welding wires, the substrate and WAAM alloys are shown in Table 1. WAAM system is constructed by Fronius CMT advanced 4000 R power supply and ABB1410 robot, as shown in Figure 1. WAAM alloy was deposited by a single pass and multi-layer forming method. The forming process is shown in Figure 2. The process parameters are as follows: wire feeding speed (WFS) 6 m\u00b7min\u22121, welding speed (TS) 10 m\u00b7ms\u22121, shielding gas flow rate 25 L\u00b7min\u22121, CONTACT Huimin Gu guhm@smm.neu.edu.cn School of Metallurgy, Northeastern University, Shenyang 110819, P. R. China \u00a9 2020 Informa UK Limited, trading as Taylor & Francis Group current 120\u2013130 A, voltage 20.4\u201321.4 V, and interlayer temperature controlled between 160\u00b0C and 180\u00b0C. The as-deposited alloys underwent solution treatment (540\u00b0C, 12 h) and ageing treatment (175\u00b0C, 4 h)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000067_metroi4.2019.8792876-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000067_metroi4.2019.8792876-Figure2-1.png",
        "caption": "Fig. 2. The original component for the antenna support.",
        "texts": [
            " MATERIALS AND METHODS The test case refers to a part, which is the support of the antenna of a satellite, in close collaboration with Thales Alenia Space as seen in Fig.1. This part started from an initial conventional design and was re-formed several times using topological optimization tools aiming to reduce the total weight of the component and maximize, in parallel, its rigidity. This process includes several steps like stress-analysis (external loads) and topological optimization (re-design) of it with respect to these specific loads. Both the original and the optimized components may be seen in Fig.2. and Fig.3, respectively. The material of this part is the Inconel 625 nickel-iron alloy which is frequently used in additive manufacturing process (PBF). The estimated weight of the initially designed part was 300 g while the final optimized design has an estimated weight of 270 g. The stress analysis and the topological optimization of the part are conducted using the Inspire SolidThinking software. Thus, using a certain procedure of iterative stress analyses and optimizations, the mechanical performance of the component is simulated and optimized"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001485_j.mechmachtheory.2019.103776-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001485_j.mechmachtheory.2019.103776-Figure5-1.png",
        "caption": "Fig. 5. 3-UU mechanism with two infinitesimal rotation mobility (a) and its directed graph (b).",
        "texts": [
            " (57) It can be seen that only the coefficient \u03b21 is left after applying Eq. (57) in the quadratic form of Eq. (31) (the corre- sponding 1st self-stress wrench basis represents a pure torque of x direction) Q = \u03b21 sin ( 2 \u03b1) ( a \u2212 b ) / ( b + a ) . (58) This quadratic form is a scalar and sign definite, if a = b, \u03b1 = 0or \u00b1 \u03c0 /2.Thus, this modified Bennett linkage is a first- order infinitesimal linkage and only possesses first-order infinitesimal mobility. 6.4. 3-UU infinitesimally mobile linkage The last example shown in Fig. 5 (a) is the Seoul National University (SNU) 3-U P U parallel mechanism with all the linear actuators locked in its home configuration (thus it is modeled as a 3-UU). The 3-U P U was thought to exhibit two extra infinitesimal rotational mobility or self-motions at its home position, this phenomenon was also attributed to a constraint singularity [45 , 46] . The radii of the fixed and moving platform are r 1 and r 2 respectively, and the height between the two bases is h . All the universal joints are anchored at the vertexes of two equilateral triangles. Each universal joint is considered as two 1-DOF revolute joints with axes intersected at right angle, so there are 12 revolute joints totally. And the 6 axes of the universal joints anchored at fixed base and the 6 axes of the universal joints anchored at moving platform form two parallel planes at current position. The directed graph of this 3-UU mechanism with two FCs is shown in Fig. 5 (b), the indexed numbers of links are omitted in the graph for convenience. The circuit matrix corresponding to the two FCs is J 1 J 2 J 3 J 4 J 5 J 6 J 7 J 8 J 9 J 10 J 11 J 12 B = ( 1 1 0 1 1 0 1 1 0 1 1 0 0 1 1 0 1 1 0 1 1 0 1 1 ) 1 2 . (59) The ordering sequences corresponding to the two FCs are 1 : 1 < 1 4 < 1 7 < 1 10 < 1 11 < 1 8 < 1 5 < 1 2 , 2 : 3 < 2 6 < 2 9 < 2 12 < 2 11 < 2 8 < 2 5 < 2 2 , (60) in which the joint J 1 and J 3 are considered as the root joints and joint J 2 as the terminal joint in each FC. The sequence matrix could be obtained accordingly by Eq. (4) but not shown here for brevity. The screw coordinates of the 12 revolute joints in current configuration in Fig. 5 (a) are \u03be1 = ( 0 \u22121 0 0 0 0 )T , \u03be2 = (\u221a 3 / 2 1 / 2 0 0 0 0 )T , \u03be3 = ( \u2212\u221a 3 / 2 1 / 2 0 0 0 0 )T , \u03be4 = ( 1 0 0 0 0 r 1 )T , \u03be5 = ( \u22121 / 2 \u221a 3 / 2 0 0 0 r 1 )T , \u03be6 = ( \u22121 / 2 \u2212\u221a 3 / 2 0 0 0 r 1 )T , \u03be7 = ( 1 0 0 0 h r 2 )T , \u03be8 = ( \u22121 / 2 \u221a 3 / 2 0 \u2212\u221a 3 h/ 2 \u2212h/ 2 r 2 )T , \u03be9 = ( \u22121 / 2 \u2212\u221a 3 / 2 0 \u221a 3 h/ 2 \u2212h/ 2 r 2 )T , \u03be10 = ( 0 \u22121 0 h 0 0 )T , \u03be11 = (\u221a 3 / 2 1 / 2 0 \u2212h/ 2 \u221a 3 h/ 2 0 )T , \u03be12 = ( \u2212\u221a 3 / 2 1 / 2 0 \u2212h/ 2 \u2212\u221a 3 h/ 2 0 )T . (61) There are two-dimensional first-order motions for this linkage with the basis matrix as N = \u239b \u239c \u239c \u239d \u2212 r 2 r 1 \u2212 r 2 r 1 0 \u221a 3 r 2 3 r 1 \u2212 \u221a 3 r 2 3 r 1 \u22122 \u221a 3 r 2 3 r 1 \u2212 \u221a 3 3 \u221a 3 3 2 \u221a 3 3 1 1 0 r 2 r 1 0 \u2212 r 2 r 1 \u221a 3 r 2 3 r 1 2 \u221a 3 r 2 3 r 1 \u221a 3 r 2 3 r 1 \u2212 \u221a 3 3 \u22122 \u221a 3 3 \u2212 \u221a 3 3 \u22121 0 1 \u239e \u239f \u239f \u23a0 T . (62) which corresponds to two infinitesimal revolute motions of the moving platform about the intersection point of three limbs and with the revolute axes parallel to the plane of platform as shown in Fig. 5 (a). The self-stress wrenches are given with two arbitrary constant \u03b21 and \u03b22 as F = ( s T 1 s T 2 )T \u03b2 = ( 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 )T ( \u03b21 \u03b22 ) . (63) The two components s 1 and s 2 represent two pure torques of z direction in corresponding FCs. The quadratic form in Eq. (39) for this two-loop linkages is Q = \u221a 3 ( r 2 1 \u2212 r 2 2 ) 3 r 2 1 ( \u03b21 Q 1 + \u03b22 Q 2 ) , (64) in which Q 1 = ( 2 \u22121 \u22121 \u22121 ) , Q 2 = ( 1 \u22122 \u22122 1 ) . (65) It could be tested that there does not exist any proper \u03b21 and \u03b22 to make Q in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001019_rpj-07-2018-0182-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001019_rpj-07-2018-0182-Figure3-1.png",
        "caption": "Figure 3 Inserting and locating the metal tube",
        "texts": [
            " Existing die assembly consists of top and bottom die sets which are made up of right and left halves. Hydroforming a Metal bellow hydroforming Prithvirajan R. et al. Rapid Prototyping Journal D ow nl oa de d by N ot tin gh am T re nt U ni ve rs ity A t 0 2: 47 3 1 M ay 2 01 9 (P T ) single convolution involves four steps and they are repeated till the required number of convolution is formed. Step 1: The preformed metal tube is inserted over the rubber bladder and positioned at a required height where the convolution is to be formed (Figure 3). Step 2: Top and bottom die halves are closed laterally to clamp the metal tube with the rubber bladder (Figure 4). Step 3: Application of hydraulic pressure inside the bladder expands the metal tube (Figure 5). This step is also referred as bulging (Kang et al., 2007). Step 4: The top die is closed vertically downwards till it touches the bottom die, the metal tube gets the \u201cU\u201d shape convolution (Figure 6). This step is also referred as folding (Kang et al., 2007). Then the formed convolution is shifted down and all the four steps are repeated till the required number of convolutions are achieved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002218_j.mechmachtheory.2019.103771-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002218_j.mechmachtheory.2019.103771-Figure13-1.png",
        "caption": "Fig. 13. Finite element analysis model of horizontal buffering mechanism.",
        "texts": [
            " Then the elastodynamic equations of the whole pitching mechanism can be derived by integrating all the element elastodynamic equations together as \u23a7 \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23a9 M p U\u0308 p + K p U p = F p \u2212 M p U\u0308 pr M p = \u2211 4 i =1 B T p i m p i B p i K p = \u2211 4 i =1 B T p i k p i B p i F p = \u2211 4 i =1 B T p i f p i (23) Likewise, the generalized force vector F p of the pitching mechanism can be obtained directly by counteracting the inter- actions between the adjacent components as F p = [ 0 0 0 0 0 0 0 0 0 F Mx F My 0 ]T (24) where F Mx and F My represent the external load acted on joint M in x- and y -directions, respectively, which can be yielded by solving the Newton\u2013Euler rigid-body dynamic equations. 3) Elastodynamic modeling of buffering mechanism The horizontal buffering mechanism extracted from the main-motion mechanism of the forging manipulator can be decomposed into five beam elements, and the motion of point M is restricted to the x- direction to ensure the output characteristics of the Hoeckens straight-line mechanism. The finite element model is shown in Fig. 13 , and the similar method is used to set up the generalized coordinates of the buffering mechanism. Similarly, subscript \u201ch\u201d is added to each parameter in the buffering mechanism to distinguish the parameters in different mechanisms. Vector U h composed of the generalized coordinates in the buffering mechanism is expressed as U h = [ U h1 U h2 \u00b7 \u00b7 \u00b7 U h15 ]T (25) Vector U\u0308 hr composed of rigid-body accelerations of each node in the buffering mechanism can be presented as U\u0308 hr = [ \u03b8\u0308PN X\u0308 N Y\u0308 N \u03b8\u0308NG X\u0308 G Y\u0308 G \u03b8\u0308NG X\u0308 M \u03b8\u0308GM \u03b8\u0308GQ X\u0308 Q Y\u0308 Q \u03b8\u0308GQ \u03b8\u0308QH \u03b8\u0308QH ]T (26) Similarly, the coordinate transmission matrix B h i is used to transform the generalized coordinates U e h i of each element into the generalized coordinates U h of the buffering mechanism, and the elastodynamic equations of each component can be obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000983_j.compgeo.2019.04.023-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000983_j.compgeo.2019.04.023-Figure4-1.png",
        "caption": "Fig. 4. Rolling resistance of a hexagon (a) and a sphere (b).",
        "texts": [
            " When representing the particle shape in simulations, more complex numerical models, especially for the contact detection, have to be developed, leading to a significantly higher degree of complexity and computational cost. One simple and effective method for accounting for the particle shape is the rolling resistance [15\u201317]. A certain rolling moment is required to roll a spherical element [42]. For this purpose, the governing equations may be modified by simply incorporating a rolling threshold. Let us consider a regular hexagon and a sphere that roll on an incline plane, as shown in Fig. 4. Stability of the hexagon with regard to rolling is ensured with its shape. For the sphere, its stability can only be ensured with a rolling moment m I 0 , and its value is limited to: m \u00b5 R pI r I I 0 (24) where RI is the radius of the sphere, and \u00b5r is the rolling friction coefficient in analogy to the sliding friction coefficient. RIneeds to be determined for a contact of two spheres with different radius. A common choice, e.g. [17], the geometric mean of the radii of two spheres is used in this study"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000108_j.biosystemseng.2019.08.002-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000108_j.biosystemseng.2019.08.002-Figure1-1.png",
        "caption": "Fig. 1 e Experimental setup: (1) Platform, (2) Supporting shaft, (3) Pitch angle adjustment device, (4) Roll angle adjustment device, (5) Base.",
        "texts": [
            " In order to increase levelling precision, and increase the vehicle operating speed range over which the levelling system works, a rapid active levelling method is proposed. As current levelling mechanisms are not completely suitable for the platform of agricultural vehicles due to their complex composition and large size, a new parallel mechanism for active levelling has been designed. The ability of the rapid active levelling method to improve levelling precision was compared with the current method. The developed agricultural vehicle levelling system was composed of a platform mechanism (to be levelled), a fourwheel vehicle chassis and a control system (Fig. 1). The size of the vehicle levelling system was 1.4 m 0.8 m 1m (L W H). Thewheelbase of the chassis was 0.792m, the wheel track was 0.765 m. The diameter of vehicle wheels was 0.275 m. The two rear wheels were driving wheels and they were equipped with a 600 W wheel motor, powered by a 48 V DC lithium ion battery. The wheels were attached to the vehicle body by a suspension system. Four infrared distance measuring sensors (Model: GP2Y0A21YK0F, Sharp Corporation, Osaka, Japan) were installed on the vehicle body in front of the 4 wheels"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000108_j.biosystemseng.2019.08.002-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000108_j.biosystemseng.2019.08.002-Figure3-1.png",
        "caption": "Fig. 3 e Simplified model of the levelling platform mechanism.",
        "texts": [
            " The levelling action started as soon as the platform tilted, thus actively preventing or reducing the inclination of the vehicle body. Because this levellingmethod provided early detection and real-time active levelling, it was designated rapid active levelling. The working procedure for the rapid active levelling method is shown in Fig. 2. As a key part of the vehicle levelling system, the platform mechanism was designed as a complete modular mechanism that could be installed on anyworking vehicle or surface using mounting holes in the base of the mechanism. A simplified model of the platform mechanism is shown in Fig. 3. As illustrated in Figs. 1 and 3, the system was composed of the platform, a supporting shaft, a base, a pitch angle adjustment device and a roll angle adjustment device. The supporting shaft consisted of cross-axle universal shaft coupling that connected the upper and lower shafts with lengths c and d, respectively. Both the pitch and roll adjustment devices had servo electric cylinders (Model: FDR044, Suzhou Fengdarui Automation Equipment Technology Co., LTD., Suzhou, Jiangsu, China) with straight shape ball joint rod end bearing where the long rod end of the straight shape ball joint rod end bearing were fixed to the platform or base and the other end was connected to the servo electric cylinders, which converted the linear motion of the servo electric cylinder into the pitch angle or roll angle change of the platform. The maximum speed of the servo electric cylinders was 200 mm s 1. When the lower surface of the platform was parallel to the upper surface of the base, the distance between the two planes being c \u00fe d. Both ends of the pitch adjustment device and the roll adjustment device were articulated, and the base was fixed to the vehicle body. As illustrated in Fig. 3, three coordinate systems were established on the levelling platform mechanism: a moving coordinate system i1j1k1 (the originwasO1) on the upper shaft, a static coordinate system i2j2k2 (the origin was O2) on the cross-axle universal shaft coupling, and a static coordinate system i3j3k3 (the origin was O3) on the base. The two coordinate systems on the upper shaft and crossaxle universal shaft coupling were coincident under the initial conditions, and the origin of coordinates was the centre of the cross-axle universal shaft coupling. The origin (O3) of the coordinate on the base was the centre of the base. In Fig. 3 P1 is the connection point between the pitch angle adjustment device and the platform. P3 is the connection point between the pitch angle adjustment device and the base, Q1 is the connection point between the roll angle adjustment device and the platform. In Fig. 3 Q3 is the connection point between the roll angle adjustment device and the base. At the initial installation, the direction of axis j1, axis j2 and axis j3 coincides with the direction of the running vehicle, and the direction of axis i1, axis i2 and axis i3 is perpendicular to the direction of running vehicle. The cross universal joint was allowed rotate around the axis i2 and axis j2 to ensure that the pitch adjustment device and the roll angle adjustment device could adjust the pitch and roll angle of the platform respectively",
            "b) where, am is the reverse adjust angle of the pitch angle adjustment device during the adjustment period, and the direction of Da is opposite to am; bm is the reverse adjust angle of the roll angle adjustment device during the adjustment period, and the direction of Db is opposite to bm. Since the two adjustment devices could only move linearly, it was necessary to calculate the displacement of the two adjustment devices according to platform adjustment angle. The controller could calculate the number of pulses and the direction ofmotion of the servomotor based on the calculated displacement, thereby controlling the two adjustment devices to complete the levelling action. As illustrated in Fig. 3, after the levelling device completed the adjustment according to the adjustment angle, the coordinates of P1, P3,Q1, andQ3 in coordinate system i3j3k3 can be obtained by transforming the coordinates twice. At the beginning of the adjustment period, the radius vector of theQ1 and P1 in coordinate system i1j1k1 (i2j2k2) can be described by Eq. (10): rp1 1 \u00bc ej1 \u00fe ck1 (10.a) rQ1 1 \u00bc ei1 \u00fe ck1 (10.b) where, \u00f0rp1 \u00de1 is the radius vector of P1 in coordinate system i1j1k1 (i2j2k2) before the platform is levelled; \u00f0rQ1 \u00de1 is the radius vector of Q1 in coordinate system i1j1k1 (i2j2k2) before the platform is levelled; The distance from P1 or Q1 to the centre of the platform is e; The distance from the platform to the centre of the cross-axle universal shaft coupling is c; The distance from the centre of the cross-axle universal shaft coupling to the base is d"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002506_biorob49111.2020.9224384-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002506_biorob49111.2020.9224384-Figure7-1.png",
        "caption": "Fig. 7. The picking action is performed in 6 steps. In the figure, green represents activated transducers involved in the picking process. (a) The manipulator is placed around the target object and brought to the center of the basin of attraction. (b) The first two levels of the transducer rings are activated and an incremented phase shift is applied to the second level of transducers, bringing the object to about 10mm off the acoustically reflective surface of the table. (c) An incremental phase shift is applied to the bottom transducer array, moving the object from 10mm to 15mm. (d) The third level ring of transducers are activated and the phase of the second and third level transducers are incremented. This moves the object from 15mm to 30mm in height. (e) The first level transducers are deactivated, and the fourth level transducers are activated. The phase of the third and fourth, transducer ring levels are incremented, moving the object from 30mm in height to 45mm in height. (f) The second level transducer ring is deactivated. The fourth level transducer ring phase is decremented. At this height, 50mm, the levitation device can be lifted from the surface of the table.",
        "texts": [
            " Additionally, the parallel jaw gripper of the PR2 robot securely fits into a groove built into the handle of the levitation device, allowing the PR2 gripper to mate with the acoustic manipulation device to optimize ease of use. This easy to grasp design allows the PR2 to pick up the device to manipulate small objects and set down the manipulator to regain the gripper\u2019s original functionality. This system functions as a practical extension to the manipulation capabilities of general purpose robotic grippers. The picking procedure for the double ring levitation device takes advantage of both phase manipulation and selective activation and deactivation. Fig. 7 shows the sequence of ring activity and phase shifts which allow the object to be translated from the surface of the table, to the center of the top pair of rings. To test the picking action of the device, the gripper was first fitted with the ultrasonic manipulation device. The picking device was then placed on the table around a polystyrene ball. The device is turned on and the ring state and phase sequence in Fig. 7 is performed manually by a human operator, allowing the ball to be picked up. Once trapped in the upper ring, the acoustic levitation device can be picked up and moved. In an ideal scenario, the robot would be capable of positioning the acoustic manipulation device with the object directly in the center of the cylindrical transducer array, however, robots designed for larger manipulation tasks, such as the PR2 used for testing, have a positioning error larger than the size of the target object. The acoustic manipulation device compensates for positioning error by increasing the \u201cgrasping\u201d range of the robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003723_1.2833803-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003723_1.2833803-Figure2-1.png",
        "caption": "Fig. 2 Schiematic of solution region",
        "texts": [
            " Furthermore, temperature distributions in the flooded configuration will dif fer little from those in the non-flooded (i.e., non-leaking) configuration, since net energy transport into the sealing re gion due to leakage is relatively small. Further, the local axial velocities in the film, primarily created by lubricant flowing around asperities, are largely unaffected by the flooded/nonflooded boundary condition. The seal is assumed axisymmetric about its centerline. A schematic diagram of tlie seal cross-section is shown in Fig. 1 and that of the solution region is shown in Fig. 2. The direction of shaft rotation is given by Journal of Tribology Copyright \u00a9 1999 by ASME JANUARY 1999, Vol. 1 2 1 / 1 Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Fig. 1 Schematic of seal cross-section the A:-coordinate, the direction along the shaft is given by the ))-coordinate, and the direction across the film is given by the z-coordinate. The shaft surface is modeled as perfectly smooth. The microgeometry of the lip is represented by deviations from the aver age lip surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002591_scems48876.2020.9352321-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002591_scems48876.2020.9352321-Figure4-1.png",
        "caption": "Fig. 4. Profile of gearbox with one planetary gear and two parallel gears. The internal structure of the WT gearbox is complicated. It is difficult to establish an accurate simulation model. According to related researches, it is known that the meshing vibration of gears has the greatest impact on the gearbox system. Therefore, we use one planetary gear combined with two parallel gears instead of the entire gearbox system in the simulation model. Fig. 5 shows the simplified model of the gearbox based on multi-stage gears. In the simplified model, each shaft is modeled as an ideal massless linear torsional spring. And the discs in the figure represent different simplified inertial bodies in the gearbox. The ring gear is assumed to be rigidly connected to the gearbox housing, and then its inertia is not considered. Meanwhile, the meshing behavior between gear pairs is simulated while ignoring the influence of gear backlash and gear meshing losses.",
        "texts": [
            " Usually, a Sigmod function is used. 1 ( ) 1 xf x e (3) Besides, the loss function of the model is defined as 21 ( ) 2 i iL y Y (4) where yi is the observed real value of the target variable, and Yi is the expectation estimated by BPNN. The most common gearbox for doubly-fed induction generators is one-stage planetary and two parallel gear 186 Authorized licensed use limited to: University of Exeter. Downloaded on May 27,2021 at 01:55:51 UTC from IEEE Xplore. Restrictions apply. structure, and its cutaway view is shown in Fig. 4. The input of the gearbox is low-speed, high torque kinetic energy from the wind wheel. According to the three-stage gear system increase speed inside the gearbox, the out of the gearbox becomes high-speed, low torque kinetic energy that meets the needs of the generator. where Jrot is the rotational inertia of the blade of the blades and hub, JPC is the carrier inertia of the planet gears, JS is the inertia of the sun gear, JGi (i=1,2,3,4) represents the parallel gears inertia. KLSS represents the low-speed shaft, which is the input shaft of the gearbox"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000492_s12206-016-0115-8-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000492_s12206-016-0115-8-Figure4-1.png",
        "caption": "Fig. 4. Line of action of the planet gear ~ ring gear mesh.",
        "texts": [],
        "surrounding_texts": [
            "motion of each element at the center points and the bearing force between the planet gear and the carrier.\n1 2 1 a a ad d b \u03b8= + \u00b4 uur uur uurr , 0 1 1 a a a\nbd d \u03b8 2 = + \u00b4\nruur uur uur (3)\n0 1 2 a a a\u03b8 \u03b8 \u03b8 .= = uur uur uur\n(4)\nIn this study, the equation of motion for each component is organized in a vector format. Mesh forces ( spF , prF ) are transferred along the unit vector in the direction of the line of action. The unit vector in the direction of the line of action was aligned along the directions from the sun gear to the planet gear, and from the planet gear to the ring gear. In the carrier and planet gear relationship, the bearing force is transferred along the horizontal, vertical unit vector ( cpn ) of piC . In addition, the vibration model of a tooth contact point defined the equivalent modulus of elasticity of the tooth contact point when the tooth profile curve of the center surface of a gear passes the 1 pitch point with consideration of the elastic deformation of a tooth [15]. The load distributed to and transferred along the tooth width was defined as the concentrated force applied to a pitch point of a crowned tooth surface, and the friction force between the tooth surfaces was neglected. The equations of motion of each component are detailed below.\n2.1 Sun gear\nFig. 2(a) illustrates the motion relationship between a sun gear and a planet gear. Eqs. (5) and (6) are the equations of motion of stations \u20181\u2019 and \u20182\u2019 of the sun gear, respectively. The mesh force spF uur from the engagement of the sun gear\nand the planet gear is applied to the sun gear. The translational\ndisplacement ( 0 1 1 ss s \u03b8\u00b4\nuuur uur ) and moment ( 2 1 1 ss s F\u00b4 uuur uur ) are addition-\nally generated from the rotational displacement ( 1 s\u03b8 uur ). In addition, the location ( 2s a uuur\n) of the mesh force applied on the sun gear body is identified so that the moment generated from the mesh force can be considered. Each vector component is shown in the Appendix.\n(5) . (6)\n2.2 Planet gear\nFigs. 2(a)-(c) illustrate the motion relationships of the planet gear and the other elements. Eqs. (7) and (8) are the equations of motion for stations \u20181\u2019 and \u20182\u2019 of the planet gear. The carrier is simultaneously engaged with the sun gear and the ring gear, and is coupled with the planet gear through bearings. Therefore, the mesh forces ( spF uur , prF uur ), bearing force ( cpF uur\n), and moment ( cpM uuuur ) from the coupling of the sun gear and the\nring gear are applied to the planet gear. The mesh force with the sun gear is tangential at body \u2018b\u2019 of the planet gear, and the mesh force with the ring gear is tangential at \u2018c\u2019. The force and moment from the bearing on the carrier are applied at the center point ( 0 0p ,c ) of each body.\n(7)\n(8)\n2.3 Ring gear\nFig. 2(b) illustrates the motion relationship between the ring gear and the planet gear. As the ring gear is engaged with the planet gear, the mesh force prF uur is applied on point \u2018d\u2019 on the\nbase circle. The additional displacement ( 0 1 r 1r r \u03b8\u00b4\nuuur uur ) and mo-\nment ( 2 1 1 rr r F\u00b4\nuuur uur ) are generated from the equation of motion\ndue to the influence of the tooth width of the ring gear as well. Eqs. (9) and (10) are the equations of motions between stations \u20181\u2019 and \u20182\u2019 of the ring gear.\n(9) . (10)\n2.4 Carrier\nFig. 2(c) illustrates the motion relationship between the planetary gear and the carrier. The carrier is coupled with the planet gear through bearings. Bearing force ( cpF uur ) is\napplied on the carrier, and an additional moment ( 2 1 1 cc c F\u00b4\nuuur uur )\nis generated at station '2' due to the width of the carrier and the force from station '1'. The bearing force ( cpF uur ) is applied\non the center point ( 0 0p ,c ) of the carrier and the planetary gear. The corresponding relationships are shown in Eqs. (11) and (12).\n(11) . (12)\n2.5 Mesh force\nThe mesh forces ( spF uur , prF uur ) between the sun gear and the\nplanet gear and between the planet gear and the ring gear are transferred through the unit vectors along the lines of action ( spn uuur , prn uuur\n), and their magnitudes ( spF , prF ) are determined by the tooth displacements ( sp\u03b4 , pr\u03b4 ) and tooth stiffness ( spk , prk ). The tooth stiffness values are the equivalent stiffnesses of the gear teeth. spe , pre are the transmission errors\n.",
            "on the lines of action. \u00b6spe ,\u00b6pre represent the magnitudes of\nthe transmission errors. In Eqs. (15) and (16), l represents the order of the harmonic component, \u03c9 represents the tooth passing frequency, sp\u03a6 , pr\u03a6 represent the phase difference\nbetween a planet gear in an arbitrary position and the reference planet gear, and rs\u03a6 represents the phase differences between the reference sun gear ~ planet gear mesh and the planet gear ~ ring gear mesh [16]. The magnitude of the transmission error is a source of vibration on the gear teeth surface, which acts as a source of vibration in a forcedvibration analysis.\n( )sp sp sp sp sp sp spF F n k \u03b4 e n= = -\nuur uuur uuur (13) ( )pr pr pr pr pr pr prF F n k \u03b4 e n= = - uur uuur uuur (14)\n\u00b6 ( )spl \u03c9t \u03a6 spl spe e e + = (15)\n\u00b6 ( )pr rsl \u03c9t \u03a6 \u03a6 prl pre e e . + + = (16)\nspn uuur , prn uuur\nare expressed as Eqs. (17) and (18), due to points \u2018a\u2019, \u2018b\u2019, \u2018c\u2019, and \u2018d\u2019, which are the tangential points of the lines of action and the base circle in Figs. 3 and 4. The line of action becomes tangential with the base circle of each gear as the line of action rotates as much as the base circle helix angle about the pitch point on the action plane. The position vector of each point can be described by the base circle radius ( cr ), the operation pressure angle ( sp pr,f f ), the tooth width (b), the base circle helix angle (\u0393) , the contact center distance ( meshC ), and the planet gear position angle (\u03c8) . Eqs. (19)- (26) show the position vectors of \u2018a\u2019, \u2018b\u2019, \u2018c\u2019, and \u2018d\u2019.\nsp abn ab =\nuuruuur uur (17)\npr cdn cd =\nuuruuur (18)\n1 1 1 1ab as s b s a s b= + = - + uur uur uur uuruur = 1 1 1 1s a s p p b- + + uur uuur uuur\n(19)\n( ) ( ) ( ) ( )1 s sp s sp s sp bs a r cos \u03b1 i r sin \u03b1 j r tan tan \u0393 k 2 \u00e6 \u00f6= + + + f\u00e7 \u00f7 \u00e8 \u00f8\nuur r r r\n(20)\n( ) ( )1 1 mesh meshs p C cos \u03c8 i C sin \u03c8 j= +\nuuur r r (21)\n( ) ( ) ( ) ( )1 p sp p sp p sp bp b r cos \u03b1 i r sin \u03b1 j r tan tan \u0393 k 2 \u00e6 \u00f6= - - + - f\u00e7 \u00f7 \u00e8 \u00f8\nuuur r r r\n(22) 1 1 1 1 1 1 1 1cd cp p d p c p d p c p r r d= + = - + = - + + uuur uuur uur uuur uur uuur uuruur (23)\n( ) ( ) ( ) ( )1 p pr p pr p pr bp c r sin \u03b1 i r cos \u03b1 j r tan tan \u0393 k 2 \u00e6 \u00f6= + + - f\u00e7 \u00f7 \u00e8 \u00f8\nuur r r r\n(24)\n( ) ( )1 1 mesh meshp r C cos \u03c8 i C sin \u03c8 j= - -\nuuur r r (25)\n( ) ( ) ( ) ( )1 r pr r pr r pr br d r sin \u03b1 i r cos \u03b1 j r tan tan \u0393 k 2 \u00e6 \u00f6= + + - f\u00e7 \u00f7 \u00e8 \u00f8\nuur r r r .\n(26)\nTooth displacement on a line of action is determined by the body displacement at each gear element, \u2018a\u2019, \u2018b\u2019, \u2018c\u2019 and \u2018d\u2019. The tooth displacement between the sun gear and the planet gear ( sp\u03b4 ) is determined by the sun gear displacement ( s\u03be ) at point \u2018a\u2019 and the planetary gear displacement ( ps\u03be ) at point \u2018b\u2019. The tooth displacement between the planetary gear and the ring gear ( pr\u03b4 ) is determined by the planetary gear displacement ( pr\u03be ) at point \u2018 c \u2019 and the ring gear displacement ( r\u03be ) at point \u2018d\u2019. Eqs. (27)-(32) express the tooth displacement of each gear element.\nsp s ps\u03b4 \u03be \u03be= - (27)\npr pr r\u03b4 \u03be \u03be= + (28)\n1 1 s s s sp\u03be d \u03b8 s a n\u00e6 \u00f6= + \u00b4 \u00d7\u00e7 \u00f7\n\u00e8 \u00f8\nuur uuruur uuur (29)\n1 1 ps p p sp\u03be d \u03b8 p b n\u00e6 \u00f6= + \u00b4 \u00d7\u00e7 \u00f7\n\u00e8 \u00f8\nuur uuuruur uuur (30)\n1 1 pr p p pr\u03be d \u03b8 p c n\u00e6 \u00f6= + \u00b4 \u00d7\u00e7 \u00f7\n\u00e8 \u00f8\nuur uuruur uuur (31)\n1 1 r r r pr\u03be d \u03b8 r d n .\u00e6 \u00f6= + \u00b4 \u00d7\u00e7 \u00f7\n\u00e8 \u00f8\nuur uuruur uuur (32)",
            "2.6 Bearing force and moment\nThe support bearing of a planet gear enables the bearing displacement to be predicted from the displacements of the planet gear and the carrier. Since the carrier and the planet gear are not coaxial, the translational displacement ( cp\u03b4 uuur ) of the bearing is computed from the translational ranges of the bearing and carrier, the rotational range of the carrier, and Eqs. (33) and (34). The bearing force due to the bearing stiffness and bearing translational displacement is applied along the unit vector ( cpn uuur ), and its magnitude ( cpF ) is computed from Eq. (35). In addition, for the support bearing of the planetary gear, a moment is generated as shown in Eqs. (37)-(40) due to the rotational displacement ( cp2\u03b4 ) of the planetary gear. Here, rk , ak and tk are the radial stiffness, axial stiffness, and tilting stiffness of the planet gear support bearing, respectively.\n1 1 1 0 1 cp c c p cp,x cp,y cp,z\u03b5 d \u03b8 c p d \u03b5 \u03b5 \u03b5= + \u00b4 - = + + uur uur uuuur uuruur uuuur uuuur uuuur\n(33)\n( )cp r cp,x cp,y a cp,z\u03b4 k \u03b5 \u03b5 k \u03b5= + + uuur uuuur uuuur uuuur\n(34)\ncp cp cpF \u03b4 n= \u00d7 uuur uuur\n(35)\ncp cp cpF F n= uur uuur\n(36)\n1 1 cp c p cp2,x cp2,y cp2,z\u03b5 \u03b8 \u03b8 \u03b5 \u03b5 \u03b5= - = + + uur uuruur uuuuur uuuuur uuuuur\n(37)\n( )cp2 t cp2,x cp2,y\u03b4 k \u03b5 \u03b5= + uuuur uuuuur uuuuur\n(38)\ncp cp2 cpM \u03b4 n= \u00d7 uuuur uuur\n(39)\n1 cp cp pM M \u03b8 .=\nuuruuuur (40)\n2.7 Planetary gear system local transfer matrix\nRearranging the equations of motion provided above yields the local transfer matrix of the planetary gear system. When the number of planet gears increases, the additional planetary gears can be considered from Eqs. (3)-(12), and can be derived using the previously employed method. The equation of motion in vector notation can be rearranged to linear algebra notation according to Table 1.\ng[T ] represents the transfer matrix of the planetary gear system, and 0g[T ] , 1g [T ] , and 2g [T ] represent the stiffness, damping, and inertial elements, respectively. g[F ] is a vari-\nable that represents the transmission error excitation force. S\u00e9 \u00f9\u00eb \u00fb is a variable that represents the translational and rotational\ndisplacements, force, and moment of each element. Variables in bold represent column variable matrices. Details of the transfer matrix elements of the planetary gear expressed in Eqs. (41)-(43) are explained in the appendix.\nn n n n n[S ] [d ,\u03b8 ,F ,M ]= (41) n 1 n n n\ng g[S ] [T ][S ] [F ]+ = + (42) 2 g 0g 1g 2g[T ] [T ] [T ]\u03bb [T \u03bb .]= + + (43)\n3. Solution technique of the transfer matrix method\n3.1 Local transfer matrix derived by the Hibner branch method\nModeling using HBM for the local transfer matrix is necessary to derive the global transfer matrix. For multi-axial systems such as external and planetary gears, the derivation must consider the coupling between each axis, but there are cases where empty spaces exist, as shown in Fig. 5. There are a total of 2 lines, 3 elements, and 4 stations. Element#3 of line#1 and element#1 of line#2 are empty spaces with no axis. Analysis of station#3 of element#2, line#1 revealed that it is equivalent to stations#3 and #4 of element#3, and the analysis of stations#1 and #2 of element#1, line#2 showed that it is equivalent to station#2 of element#2. When the state variables of lines#1 and #2 of station#1 are 1[ 1 ]S and 1[ 2 ]S , respectively,\nthe total state variable of station#1 is 1[ ] =S 1 1[[ 1 ],[ 2 ],1]TS S . Line#1 of element number# 1 in Fig. 5 is empty, but when the state variable of station#2 of line#2 is made equivalent up to station#1, local transfer matrix modeling of element number 1 is possible even when there is empty space. Because of this, modeling using the transfer matrix method is possible for complex, multi-axial gear systems.\n2 1 1\u00e9 \u00f9 \u00e9 \u00f9 \u00e9 \u00f9=\u00eb \u00fb \u00eb \u00fb \u00eb \u00fbS T S (44)\n2 1\n2 1 1 10 0 2 0 0 2 . 1 10 0 1 \u00e9 \u00f9 \u00e9 \u00f9\u00e9 \u00f9\u00e9 \u00f9 \u00e9 \u00f9 \u00e9 \u00f9\u00eb \u00fb \u00eb \u00fb \u00eb \u00fb\u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa=> = \u00e9 \u00f9 \u00e9 \u00f9 \u00e9 \u00f9\u00eb \u00fb \u00eb \u00fb \u00eb \u00fb\u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa\u00e9 \u00f9 \u00e9 \u00f9\u00eb \u00fb \u00eb \u00fb\u00eb \u00fb\u00eb \u00fb \u00eb \u00fb T T S ST S I S"
        ]
    },
    {
        "image_filename": "designv11_14_0002182_302-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002182_302-Figure2-1.png",
        "caption": "Fig. 2. Diagram of cylindrical spool",
        "texts": [
            " The introduction of moving axes then enables the equations of motion to be written in a simpler form, admitting of exact solution. By comparing solutions for different positions of D, it is then possible to obtain an estimate of the variations of tension occurring in the real case. 2. T H E E Q U A T I O N S O F M O T I O N Consider the steady state case in which inextensible thread is pulled with constant speed V from a spool in the form of a right circular cylinder of radius a and height Z, (Fig. 2). The unwound portion of thread CD, which is assumed to be in contact with the spool at all points, lies on a curve of constant shape, which moves round the spool with angular velocity Vja. At the same time the thread is moving along the curve DC with velocity V, and the frictional force at any point on curve DC therefore acts in the opposite direction to the resultant of these two velocities. BRITISH JOURNAL OF APPLIED PHYSICS 142 VOL. a. APRIL 1957 Vaviations in tension of an unwinding thread It is convenient to use cylindrical polar co-ordinates, (",
            "V2 - To) (6) Eliminating N and a\u2018\u2018 from equations (l), (2) and ( 5 ) : \u2019T = (TjmV2) - 1 Thus, from equation (6), c j m V K a8\u2019 = - __ - 1 = - - 1 where K = - cjmV2 (8j - T 2nd from equations (4) and (7) we obtain finally: where a = 4(2K + 3 -+ d ( 4 K -k 9) p = +[2K 1. 3 - d ( 4 K +- 9)] Integrating with limits 0 and Zi for Z and T D , T~ for T we obtain: 3 . T H E L I M I T S O F I N T E G R A T I O N Suppose, now, that the lower limit of integration in equation ( 5 ) is taken at D. Then, since the thread is wound perpendicular to the axis of the spool 8; = l/a. Thus, from equation (8), 2rD = K i.e. \u2018 T ~ = K/2 where T D is the value of T at D. If the thread leaves the cylinder at an angle $ with the horizontal, as shown in Fig. 2, a& = cos 9 and \u2019ic = K/(1 - cos 4) = hD/ (1 J- cos $) The limits of integration in equation (9) can therefore be written as: T~ and 27-,/(1 T cos $) and equation (9) therefore becomes: 4 . T H E P O S S I B L E R A N G E OF S O L U T I O N There appears to be little information available on the values of r occurring in practice, but it is possible to impose certain theoretical limits. Thus, since T/m V 2 is essentially positive, T > - 1 and K > - 2. Consider, now, the expression for the normal reaction, N, when 8\u2018 is eliminated between equations (1) and (8)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001315_tiv.2019.2955904-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001315_tiv.2019.2955904-Figure6-1.png",
        "caption": "Fig. 6. Schematic of the top view of the vehicle, where \u03c6 is the steering angle, \u03c1 is the turning radius, and t and n are the unit tangent and normal vectors of the path at the midpoint of the rear axle, respectively.",
        "texts": [
            " 4, the clearance C between two vehicles can be obtained as: C = 2M cos\u03b2 \u2212 2\u039b sin\u03b2 , (8) which becomes zero when M = \u039b tan\u03b2 . (9) Based on the path c(\u03b6), the motor/engine-torque limitation of the vehicle, and the equation of motion, the time-optimal vehicle control problem can be solved using the algorithm by Bobrow et al. [3] which is studied in Sections V and VI. We define a 3D Cartesian reference coordinate system oxyz, as depicted in Fig. 5. The unit vector along the positive z-axis is given as k. The model of our vehicle is illustrated in Fig. 6. The front wheels are directional, while the rear wheels are fixed. The planar path of the vehicle is determined from the position (x, y) of the midpoint of the rear axle [1]. The unit tangent t, normal n, and binormal b vectors define a moving orthogonal frame along the path c(\u03b6), called the Frenet-Serret frame, where (t, n, b) are computed as follows: t = c\u2032 |c\u2032| , b = c\u2032 \u00d7 c\u2032\u2032 |c\u2032 \u00d7 c\u2032\u2032| , n = b\u00d7 t. (10) The Darboux frame (t, N, q) is a similar orthogonal frame, including the surface geometry, evaluated as t = Ruu \u2032 + Rvv \u2032 |Ruu\u2032 + Rvv\u2032| , N = Ru \u00d7Rv |Ru \u00d7Rv| , q = N\u00d7 t"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002218_j.mechmachtheory.2019.103771-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002218_j.mechmachtheory.2019.103771-Figure6-1.png",
        "caption": "Fig. 6. Virtual prototype model of the novel forging manipulator.",
        "texts": [
            " With a carrying capacity of 2t/5t m, the geometric parameters of this novel forging manipulator are obtained through dimensional optimization, as shown in Table 1 . For the convenience of describing the motions of the manipulator, a Cartesian coordinate system A-xy is attached to point A to easily obtain the positions of other points in the same coordinate system, which are B ( \u22121100, 170), P ( \u2212600, \u2212530), H ( \u2212600, \u2212930), I ( \u2212800, \u22121330), respectively. After the structural design and optimization, a virtual prototype model of the 2t/5t m forging manipulator is constructed, as shown in Fig. 6 . The mass and the moment of inertia of each component can be obtained directly by measuring the 3D model, and the specific values are listed in Table 2 . The dynamic analysis includes two main aspects: rigid-body dynamics and elastodynamics. The elastodynamic analysis is based on some analytical results of the rigid-body dynamics, such as accelerations of rigid-body movements and the constraint forces of each component, thus investigation of the rigid-body dynamics is necessary before the elastodynamic analysis for the novel forging manipulator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000930_s11044-019-09676-2-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000930_s11044-019-09676-2-Figure3-1.png",
        "caption": "Fig. 3 Function g(\u03be), see Eq. (65)",
        "texts": [
            " Similarly to (44), we have T 2 = \u03c12 2 (\u03be/ cos \u03be)2. (63) Let us now insert formulas (60) and (61) into Eq. (32). We obtain, similarly to (45), the following equation: z1 \u2212 z0 = \u03c12 2g(\u03be)/4 (64) where the function g(\u03be) = [ 2\u03be + sin(2\u03be) ] / cos2 \u03be (65) is introduced. Function g(\u03be) is odd and monotone on the interval |\u03be | < \u03c0/2, it grows from \u2212\u221e to \u221e as \u03be changes from \u2212\u03c0/2 to \u03c0/2. Its properties are summarized below: g(\u2212\u03be) = \u2212g(\u03be), g(\u03be) \u223c 4\u03be as \u03be \u2192 0, g(\u03be) = \u00b1\u03c0\u03b2\u22122, \u03b2 = \u03c0/2 \u2213 \u03be > 0 as \u03be \u2192 \u00b1\u03c0/2. (66) The graph of function g(\u03be) is shown in Fig. 3. Let us, similarly as in Sect. 6, describe the procedure that gives the solution of Problem 2. If z1 = z0, then we satisfy Eq. (64) by \u03be = 0. Here, we have again the trivial case where the terminal conditions are fulfilled already at the initial state, and T = 0. We will consider below the case where z1 = z0 and suppose first that \u03c12 > 0. Then Eq. (64) for \u03be , by virtue of properties (66) of function g(\u03be), has a unique solution. According to relationships (65) and (66), we have sign \u03be = sign ( z1 \u2212 z0 ) , |\u03be | < \u03c0/2, cos \u03be > 0 for this solution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000038_j.apsusc.2019.143594-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000038_j.apsusc.2019.143594-Figure3-1.png",
        "caption": "Fig. 3. Three-dimensional geometry used in numerical analysis, boundary conditions of a radiative heating source, and consequent heat transfer through the target powder bed.",
        "texts": [
            " The impact of the object functions, the optimal power and exposure time of radiative heating source, can be determined within a confined design domain. We focus on the thermal conditions required for melting a target area and the critical limit to prevent deterioration in performance. The performance deteriorates when the molten area expands beyond the intended linewidth due to over-melting. This is primarily caused by the excess heat dissipation and Marangoni effects, which accompany the anisotropic expansion of the molten pool. Further details will be provided in the Results and Discussion section. Fig. 3 shows the geometric domain used in our numerical analysis to investigate the heat transfer characteristics of the laser melting process. Our target structure is a standalone workpiece. We investigate the variations in transient heat transfer conditions as the system is exposed to the laser. We focus on the effects of the power and exposure time on stationary heating and subsequent cooling behaviors. These should be evaluated independently in advance so that we can then investigate parameters relating to the movements of the laser, such as velocity, hatch style, and layer thickness",
            " The laser beam has an effective Gaussian beam power distribution with a symmetrical irradiance distribution. We assume that the laser irradiance is symmetric about the direction of propagation and, in most cases, the maximum irradiance of the power per unit area I0 is at the center of the beam pattern. The beam irradiance of the fundamental mode is defined as I(r)= I0e\u22122r2/r02, where r and r0 are the radial distance from the center and the beam radius corresponding to the point where the irradiance diminishes to 1/e2, respectively [38]. Herein, r0 is 35 \u03bcm. The inset in Fig. 3 describes the irradiance profile on the heating surface. Taking into account this radial distribution of the irradiance, we can use the expression for the input power of the laser to define the surface heat flux, ql\u2033, as follows [38]: \u2032 = \u2032 \u2212q P \u03c0r e2 l l r r 0 2 2 /2 0 2 (4) where Pl is the laser power. We use a conventional ytterbium (Yb) laser with a wavelength of 1070 nm and powers of 50, 100 and 200W, which are typical for laser-induced manufacturing [9]. They are equivalent to apparent surface heat flux of 1",
            " Under locally concentrated heating with moderate secondary heat dissipation through convection and radiation on the exterior surfaces, conduction is the principal process governing both the local and the overall heat transfer characteristics. Convection effects in the chamber should be minimized to prevent the material losses that can arise due to blowing and oxidation of the material. As this process is dominated by conduction, heat dissipation or transfer in a solid media is explained by the Fourier's law of conduction (Eq. (2)). Based on the relationship, we investigate local and overall heat transfer characteristics analytically. In a powder bed system (shown in Fig. 3), these characteristics can be determined by the effective properties of the powder bed, the input heating power, and the exposure time [8,20,25]. The absorbed energy can be quantified in terms of these parameters. Herein, we consider macroscopic approaches based on the plausible assumption that powder beds are homogeneous in confined domains [24,38]. When the target powder bed has high thermal conductivity, this will dominate the thermal response of the powder. Conduction effects within the powder bed lead to the incident thermal energy propagating towards the surrounding powder bed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure6-1.png",
        "caption": "Fig. 6. Rotation by b = p=6, followed by a rotation by a = p=6.",
        "texts": [],
        "surrounding_texts": [
            "The objective of our method is to estimate all of the unique components of I and rC=B. The strategy relies on being able to translate the object by rB and then rotate it by R with respect to frame O which is defined by the apparatus. The term IR in Eq. (15) can be expanded to: IR \u00bc R Ixx \u00fem yC 2 \u00fe z2C Ixy mxCyC Ixz mxCzC Ixy mxCyC Iyy \u00fem xC2 \u00fe z2C Iyz myCzC Ixz mxCzC Iyz myCzC Izz \u00fem xC2 \u00fe y2C 0 B@ 1 CART \u00f016\u00de The selection of R is very important. It was found that the rotations presented in Figs. 5 and 6 are of mathematical and experimental convenience. For experimental testing, the object is first rotated about yO by angle b, then about x1 by angle a, as shown in Figures 5 and 6, where the bases x1; y1; z1 and x2; y2; z2 are the result of the first and second rotations respectively. The resulting rotation matrix is written as: R \u00bc Ry;bRx;a \u00bc cos b\u00f0 \u00de sin a\u00f0 \u00de sin b\u00f0 \u00de cos a\u00f0 \u00de sin b\u00f0 \u00de 0 cos a\u00f0 \u00de sin a\u00f0 \u00de sin b\u00f0 \u00de cos b\u00f0 \u00de sin a\u00f0 \u00de cos a\u00f0 \u00de cos b\u00f0 \u00de 0 B@ 1 CA \u00f017\u00de We focus on the bottom-right component of IR(15), denoted as IRzz, whose value can be measured directly (c.f. Section 4): i 1 2 3 IRzz \u00bc Ixx sin 2 b\u00f0 \u00de 2Ixy sin a\u00f0 \u00de sin b\u00f0 \u00de cos b\u00f0 \u00de \u00fe 2 cos a\u00f0 \u00de cos b\u00f0 \u00de sin b\u00f0 \u00de mxCzC Ixz\u00f0 \u00de\u00f0 \u00fe sin a\u00f0 \u00de cos b\u00f0 \u00de Iyz myCzC \u00de \u00fe sin2 a\u00f0 \u00de cos2 b\u00f0 \u00de Iyy \u00fem xC2 \u00fe z2C \u00fe cos2 a\u00f0 \u00de cos2 b\u00f0 \u00de Izz \u00fem xC2 \u00fe y2C \u00femxCyC sin a\u00f0 \u00de sin 2b\u00f0 \u00de \u00femy2C sin 2 b\u00f0 \u00de \u00femz2C sin 2 b\u00f0 \u00de \u00f018\u00de where from (13), xC \u00bc xB \u00fe xC=B \u00f019\u00de yC \u00bc yB \u00fe yC=B \u00f020\u00de zC \u00bc zB \u00fe zC=B \u00f021\u00de The choice of rotation and translation configuration parameters b;a; xB; yB, and zB can simplify the experimental determination of the desired inertia tensor I and CM location vector rC=B. This method requires a minimum of nine experimental configurations to estimate nine unknowns: Ixx; Iyy; Izz; Ixy; Ixz; Iyz; xC=B; yC=B; zC=B. It is assumed that the mass of the object, m, is known or can be measured. Motion measurements allow estimation of IOzz, the total mass moment of inertia about the oscillation axis, which is the summation of the inertia of the rotating assembly (plate, attachment blocks, etc) IP and the inertia of the measured object IRzz. For a given experiment with index i, the inertia corresponding to the measured object can be found from: IRzz;i \u00bc IOzz;i IP;i :\u00bc Ie;i \u00f022\u00de The terms IP;i can be estimated through direct experiments, which will be done in Section 4.7. It is worth noting that IOzz;i is assumed to be constant during a given experiment, which means the parameters b;a; xB; yB, and zB are configured before the tests, and remain constant."
        ]
    },
    {
        "image_filename": "designv11_14_0003727_1.2831176-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003727_1.2831176-Figure2-1.png",
        "caption": "Fig. 2 Approximated boundary of SA(t) ment(s) of the workpiece geometry",
        "texts": [
            " Linear path workpiece geometry essentially requires a two-dimensional reg ularized Boolean (difference) operation. This operation basi cally consists of examining the containment and intersections between the boundaries of the area swept by the cutter and the workpiece, and updating the workpiece geometry. The updating procedure consists of creation of new vertices/edges/faces and/ or modification of existing workpiece boundary elements. The boundary of the area swept by the cutter in time interval At, SA(t), is approximated as shown in Fig. 2. Based on the type of cutter feed path, two basic approximations of SA(t) are possible. For a linear path, SA(t) consists of four elements: two semi-circular segments, Efc(t) and Ebc(t) at the front and back of the cutter, and two line segments, Es!(t) and Esr(0 on either side of cutter. Both Efc(t) and Ehc(t) have the same radius of Rc and center at Oc(t - At) and Oc(t), respectively. Similarly, for a circular path, SA(t) will also consist of four edges when Rf(j) is greater than Rc (Fig. 2b). However, the linear Es,(t) and Esr(t) in the previous case are replaced here by circular Esl(t) and Es,.(t) with radii Rfij) + Rc and Rfij) \u2014 Rc, respec tively. Both arcs are centered at the center of path segment (j). When R/ij) < Rc, either edge E\u201eit) or Es,it) will not present, depending on whether the circular path is in the CW or CCW direction. Intersections between the edge elements of SAit) and the boundary of the workpiece at (f - At), given by Gpit \u2014 At), are computed. If a single valid intersection point is found be tween an edge of Gpit \u2014 At) and the boundary of SAit), as shown in Fig. 2a, the vertex of the edge contained in SAit) is modified to point to the newly found intersection point. Two valid intersections between an edge of Gpit - A;) and the boundary of SA (t) can occur (see Fig. 2b): (i) with both vertices of the edge/arc outside the boundary of SA(t), and (ii) with both vertices of the edge /arc within the boundary of SA(t). In Journal of Manufacturing Science and Engineering the first case, the edge/arc of the workpiece is split into two separate edges having the intersection points as their end and start points, respectively. In the latter case, the end points are modified to two intersection points. If an edge/arc of the workpiece is contained entirely in SAit), that edge/arc is removed from the B-rep of the workpiece"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure9-1.png",
        "caption": "Fig. 9 Schematic diagram of working rake angle",
        "texts": [
            " For convenience, the cutter with curved rake face is named cutter A, and the cutter with plane rake face is named cutter B. It is obvious that the working rake angle of cutter A is just equal to \u03b3o since the rake face takes the working rake angle as the design variable directly. However, the working rake angle of cutter B is not distinct since its rake face is designed according to the shape rake angle \u03b3e. For comparison, the working rake angle of cutter B is calculated by use of the reference system as shown in Fig. 9. The meaning of each reference plane is the same as that in Section 2.2. Furthermore, tp represents the direction vector of the intersection curve of plane Po and rake face Pj, and ve21 represents the unit vector of v21. Since vector tp locates in plane Po which is defined by N1 and v21, andN1 is perpendicular to v21, the following equation can be obtained. tp \u00bc cos\u03b3pN1\u2212sin\u03b3pv e 21 \u2264u \u00f016\u00de where \u03b3p represents the working rake angle of cutter B. Since tp also locates in the rake face Pj, and np is the normal vector of Pj, the following equation can be obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002721_j.aej.2021.01.012-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002721_j.aej.2021.01.012-Figure8-1.png",
        "caption": "Fig. 8 Generation of worm wheel tooth flank surface: a) flank surface of a machining worm model, b) surface generated by the tool external edge, c) obtained region II, d) obtained regions I and III from surface (b), e) worm wheel tooth flank surface.",
        "texts": [
            " The normal vector in the case of the rectilinear axial profile of the worm hob is described as n 2 0\u00f0 \u00de 1 \u00bc n 2 0\u00f0 \u00de x1 n 2 0\u00f0 \u00de y1 n 2 0\u00f0 \u00de z1 2 66664 3 77775 \u00bc L 2 0 1 0 \u00f0@r 1 0\u00f0 \u00de 1 @u1 @r 1 0\u00f0 \u00de 1 @u \u00de \u00f011\u00de where L 2 0 1 0 is the transformation matrix from system 1 0 to 2 0 , which is obtained by deleting the last line and the last column of the homogeneous matrix of equation (10), and this transformation matrix can be shown as L 2 0 1 0 \u00bc cos\u00f0u0 1\u00de sin\u00f0u0 1\u00de 0 cos\u00f0u0 2\u00de sin\u00f0u 0 1\u00de cos u 0 2 cos u 0 1 sin\u00f0u0 1\u00de sin u 0 1 sin\u00f0u0 2\u00de cos\u00f0u0 1\u00desin u 0 2 cos\u00f0u0 2\u00de 2 64 3 75 \u00f012\u00de The normal vector of the surface in the case of the arc axial profile of globoid worm is provided by n 2 0\u00f0 \u00de 1 \u00bc n 2 0\u00f0 \u00de x1 n 2 0\u00f0 \u00de y1 n 2 0\u00f0 \u00de z1 2 66664 3 77775 \u00bc L 2 0 1 0 \u00f0@r 1 0\u00f0 \u00de 1 @u1 @r 1 0\u00f0 \u00de 1 @h \u00de \u00f013\u00de The tangential vector is calculated based on the kinematics of worm wheel machining and is expressed as v 2 0\u00f0 \u00de 1 \u00bc v 2\u00f0 \u00de x1 v 2 0\u00f0 \u00de y1 v 2 0\u00f0 \u00de z1 2 6664 3 7775 \u00bc dr \u00f020 \u00de 1 du0 2 \u00bc dM 0 2 0 1 0 du0 2 r \u00f010 \u00de 1 \u00f014\u00de After solving equation (8), a set of solutions u1 for given values of parameter u (for the rectilinear profile) or a set of solutions u1 for given values of parameter h (in the case of the arc profile), is obtained. These parameters determine where the linear contact of the worm hob and worm wheel occurs [4]. If they are inputted into the tool surface equation, one gets the contact lines presented in the tool x 0 1y 0 1z 0 1 system. r \u00f010 \u00de c \u00bc r 1 0\u00f0 \u00de 1 u1; u\u00f0 \u00de \u00f0rectilinear profile\u00de r \u00f010 \u00de c \u00bc r 1 0\u00f0 \u00de 1 u1; h\u00f0 \u00de \u00f0arc profile\u00de \u00f015\u00de The contact lines for the tool of a particular profile are presented in Fig. 7. The generation of the worm wheel tooth flank surface is demonstrated in Fig. 8. In the considerations concerning the determination of the worm wheel mathematical model, the worm model is turned in relation to the z 0 1 axis so that the extreme cutting edge of the tool is located in the central plane, that is, plane y1z1 \u00bc y2z2 (Fig. 1). This position of the worm model is called the basic position. This operation simplifies the presentation of the mathematical model of the worm wheel tooth flank. The parameter u 0 1b is also introduced. It is the angle by which the extreme cutting edge of the machining worm will coincide with the central plane",
            " The contact line from equation (16) is transformed to the defined position: r 1 0\u00f0 \u00de 2II \u00f0i; j\u00de \u00bc M12 M0 22 0 M21 r 1 0\u00f0 \u00de c \u00f017\u00de In matrix M 0 2 0 2 of equation (17), u 0 2 \u00bc Du 0 1 i is inserted. The set of determined and transformed contact lines represents region II of the worm wheel tooth flank (r 1 0\u00f0 \u00de 2II ). The designation (i, j) by the position vector defines the numerical representation of the surface and refers to the indices of the solutions table (i and j are natural numbers). An exemplary solution for region II is presented in Fig. 8c. Regions I and III result from shaping the worm wheel tooth flank by the tool external cutting edge [4,12,13]. It is necessary to make the transformation of the tool axial profile r 1 0\u00f0 \u00de 1 u1\u00bcu1p\u00f0 \u00de using equation (5) (rectilinear profile) or (7) (arc profile) and then transform it to the central plane: r 1 0\u00f0 \u00de 1 u1\u00bcu1p\u00f0 \u00de \u00bc M 0 11 0 r 1 0\u00f0 \u00de 1 u1\u00bcu1p\u00f0 \u00de \u00f018\u00de In matrix M 0 11 0 of equation (18), u 0 1 \u00bc u 0 1b is inserted. The surface generated during processing by the tool external edge and presented in the processing worm system is expressed by equation (19): r 1 0\u00f0 \u00de 2 \u00bc M 1 0 1 M12 M2 0 2 M21 r 1 0\u00f0 \u00de 1 u1\u00bcu1p\u00f0 \u00de \u00f019\u00de In matrix M 1 0 1 of equation (19), it is necessary to select the range of parameter u1 to obtain the surface of the worm wheel tooth flank of a given face width. This is done using the values of u1p and u1k. In Fig. 8b, the surface generated during processing by the tool external edge and determined based on equation (19) is presented. It is necessary to separate regions I and III from the surface shown in Fig. 8b. The boundaries of the zones are the contact lines lying in the area of the external edge of the cutting tool (Fig. 7) [19]. For region I, it is the contact line that does not lie in the tool axial plane (r \u00f010 \u00de c2 \u00bc r \u00f010 \u00de c i; 2\u00f0 \u00de) (Fig. 7). The separated zones I and III of the worm wheel tooth flank surface generated during processing by the tool external edge are illustrated in Fig. 8d. The algorithm for selecting region I (r 1 0\u00f0 \u00de 2I \u00de consists of checking the condition r 1 0\u00f0 \u00de 2x 1 0 i; j\u00f0 \u00de < r \u00f010 \u00de c1x10 i; 1\u00f0 \u00de \u00f020\u00de where r 1 0\u00f0 \u00de 2x 1 0 i; j\u00f0 \u00de is the element from the table of coordinates x 1 0 of the worm wheel side surface r 1 0\u00f0 \u00de 2 determined based on (19) and generated during machining by the tool extreme edge, r \u00f010 \u00de c1x10 i; 1\u00f0 \u00de is the element of the table of coordinates x 1 0 of the contact line r \u00f010 \u00de c1 (Fig. 7), and i; j are natural numbers. Equation (20) defines the range of coordinates i; j\u00f0 \u00de of the table for region I",
            " It consists of verifying the condition r 1 0\u00f0 \u00de 2x 1 0 i; j\u00f0 \u00de > r \u00f010 \u00de c2x10 i; 2\u00f0 \u00de \u00f022\u00de where r \u00f010 \u00de c2x10 i; 2\u00f0 \u00de is the element of the table of coordinates x 1 0 of the contact line r \u00f010 \u00de c2 (Fig. 7). Equation (22) specifies the range of coordinates i; j\u00f0 \u00de of the table for region III. Thus, it can be expressed as r 1 0\u00f0 \u00de 2III i; j\u00f0 \u00de \u00bc r 1 0\u00f0 \u00de 2 i; j\u00f0 \u00de \u00f023\u00de where coordinates (i; j) meet the condition of (22). Regions I, II, and III of the worm wheel flank surface have to be joined (Fig. 8e): r 1 0\u00f0 \u00de 2 \u00bc r 1 0\u00f0 \u00de 2I [ r 1 0\u00f0 \u00de 2II [ r 1 0\u00f0 \u00de 2III \u00f024\u00de The joining consists of assigning to vector r 1 0\u00f0 \u00de 2 i; j\u00f0 \u00de the elements of particular regions and reindexing the elements. The process of determining regions I, II, and III of the worm wheel generated by the tool of arc profile is analogous. The obtained worm wheel tooth flank surface, shaped by the tool of concave and convex machining worm, is depicted in Fig. 9. The algorithms for determining the worm wheel tooth flanks of globoid worm gear of rectilinear and arc axial profiles are similar"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001946_icra40945.2020.9197517-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001946_icra40945.2020.9197517-Figure3-1.png",
        "caption": "Fig. 3. When a virtual plane is created at (a) the left hand, (b) the right hand, or (c) the midpoint between the two hands, another virtual plane with the same orientation is created at the center of the end effector of the robot arm. The handheld plane acts as a motion proxy to the one at the end effector.",
        "texts": [
            " Our technique allows fluid clutching with one, two, and between hands for high dexterity and comfort. Initially, the user holds two low-cost, lightweight 6-DOF motion-tracked controllers. By activating the controllers, the user can create and hold a virtual plane (Fig. 2) with one hand (e.g. as if holding a small plate in Fig. 2c, f) or two hands (e.g. as if holding a large tray in Fig. 2i). At the same time, another virtual plane with the same orientation is created at the position of the end effector of the robot (Fig. 3). When the handheld plane is translated and rotated, the motion deltas are transmitted to the plane at the end effector in real time (Fig. 4). As a result, the end effector follows hand motions as if the user is holding and moving a plane that is physically attached to the end effector. When the controllers are deactivated, the virtual planes are destroyed and the motion deltas are no longer transmitted, so the user can freely walk to a better standpoint or take a more comfortable posture without affecting the robot during transit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002220_j.triboint.2020.106167-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002220_j.triboint.2020.106167-Figure2-1.png",
        "caption": "Fig. 2. Experimental setup.",
        "texts": [
            " The size of the analyte and setup influence the spatial resolution by increasing the distance between the radioactive source and the analyte; the spatial resolution (length unit, e.g., nm) in an experiment is theoretically approximately proportional to the distance. With of all these factors considered, the actual spatial resolution achieved in this work was 2 \u03bcm. During contact mechanics experiments, pressure must be applied and maintained to the contact interfaces. An experimental apparatus was designed to achieve this (Fig. 2). The setup consisted of a regulating bolt, transparent tube, pencil rod, pressure sensor, and load bolt, among other components. The samples were confined to the transparent tube and assembled between the regulating bolt and pencil rod, and the regulating bolt was used to align the sample to the measurement region. The squeeze load was applied by screwing the load bolt. After measurement by the sensor, the load was transmitted to the samples and finally squeezed on the rough surfaces. The tube was designed to be slender to reduce the size of the setup; this helped achieve higher-resolution CT measurements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003715_s0301-679x(98)00106-6-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003715_s0301-679x(98)00106-6-Figure5-1.png",
        "caption": "Fig. 5. Photoelastic reflection test rig.",
        "texts": [
            " The test rig reproduces the real working conditions of the seal when mounted in the housing. To determine the seal stress and strain field and to observe the cross section, a segment has been cut from the ring seal for testing as a plane specimen. This deter- mines a different behaviour between the seal segment and the real seal. The cross section specimen in the test rig works in a plane stress field but the error introduced is small because the ratio of overall seal ring radius to the radial thickness is large [4,13]. The test rig is represented in Fig. 5. It is made of a base on which the test chamber is fixed. The seal segment is mounted in the test chamber. This chamber is pressurised by compressed air. The pneumatic cylinder drags the plate and the motion is driven by the linear bearings coupled with special guiding rods fixed to the base. The plate represents the cylinder rod with reciprocating motion. The plate material and roughness are the same as a cylinder rod as in the real working conditions. In Fig. 6 a scheme of the test chamber for the photoel- astic measurements shows the window, made of a resistant safety glass, and the seal segment cross section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000729_icacci.2016.7732065-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000729_icacci.2016.7732065-Figure1-1.png",
        "caption": "Fig. 1. Mobie robot Kinematics model",
        "texts": [
            " But sometimes this method fails in generating a path that reaches a goal location. Fuzzy logic is used to improve the POP algorithm by generationg smoother paths from the begining point to the end point. Another algorithm called Ant Colony Optimization (ACO) is used for mobile robot navigation in [12] where the mobile robot\u2019s optimized motion without collision with obstacles is presented. III. METHODOLOGY The mobile robot\u2019s kinematics architecture is same as specified in [13]. The robot model is shown in figure1 with VLeft and VRight as linear velocities. Left is the left wheel\u2019s angular velocity and Right is the right wheel\u2019s angularvelocity. The linear velocities of left and right wheels are calculated from the corresponding angular velocities using equation (1) . VLeft = r Left and VRight = r Right (1) Were r is the wheel radius of the two wheels and V Robot is the mobile robot\u2019s linear velocity. VRobot = (VRight + VLeft ) /2 (2) Robot = (VRight \u2013 VLeft ) / D = (r Left - r Right) / D (3) The dynamic model of the robot is given as matrix form in equation(4) that relates VRobot, Robot and (orientation of robot)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001525_acsami.9b23195-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001525_acsami.9b23195-Figure1-1.png",
        "caption": "Figure 1. Schematic illustration of the photopolymerization of a hydrogel precursor, self-rolling of the hydrogel sheet, and shape deformation of the nanocomposite hydrogel tube under an NIR light irradiation.",
        "texts": [
            " of the membrane when the constraint of the strain was removed. However, most of the rolled-up tubes reported before are lacking in their capability for remote control, which can be beneficial for the light-controlled release of cell/drug.57 In addition, the realization of self-rolling through the fabrication of a layered sheet is a relatively tedious process and is desired to be improved. Herein we present an easy and controllable approach to fabricate self-rolling nanocomposite hydrogel tubes and study their light-guided shape deformations (Figure 1). Magnetic nanorods were introduced in the poly(N-isopropylacrylamide) hydrogel precursor, which was then subjected to an ultraviolet (UV) polymerization. Due to the UV light intensity decay resulting from the UV light extinction of magnetic nanorods, a strain gradient was generated across the thickness of the formed hydrogel sheet during the photopolymerization process. After the removal of the strain constraint from the hydrogel fabrication mold, the nanocomposite sheet deformed spontaneously"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003066_j.compeleceng.2021.107267-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003066_j.compeleceng.2021.107267-Figure11-1.png",
        "caption": "Fig. 11. Basic free body diagram showing the basic forces in the rotors that allow to stabilize the testbed.",
        "texts": [
            " Quadcopters are part of the Unmanned Aerial Vehicles (UAVs), which (as discussed in [28]) have gained greater popularity in the last decade due to the fast-increasing technological advances in Unmanned Aircraft Equipment (UAE). Therefore, developing and controlling Quadcopters, represents a leading-edge topic that excites students to learn about. Thus, summarizing step-by-step the proposed V-Model structure for teaching fuzzy controllers, and following the model from Fig. 5: 1. Prototype and Fuzzy controller concepts Introduction: In this stage, all the mathematical theory of the system modeling is introduced to the students (as exemplified in Fig. 11), which brings out the control objectives in the system; allowing to also introduce the fuzzy logic main concepts (as exemplified in Fig. 9) and their implementation into control systems. 2. Prototype requirements definition: After presenting the theoretical framework of the fuzzy controller for the quadcopter, the prototype requirements for the system to be designed and implemented by the student is presented, where the main objective at this point is giving them a clear goal of what the expected results are"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000506_s11771-016-3101-5-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000506_s11771-016-3101-5-Figure4-1.png",
        "caption": "Fig. 4 Coordinate systems applied for pinion generation",
        "texts": [
            " (3), the unit normal to the pinion generating surface ( ) p a\u03a3 is determined as ( ) ( ) ( ) ( ) 2 2 1 1 1 1 1 1 2 2 01 1 1 1 1 1 1 1 1 1 2 2 1 1 1 1 1 cos sin sin 1, cos sin cos sin cos R s s s R s s R R s s \u03b1 \u03b1 \u03b8 \u03b8 \u03b1 \u03b1 \u03b8 \u03b1 \u03b1  \u2212 \u2212 \u22c5     = \u2212 \u2212 \u22c5     \u00b1 \u2212 + \u22c5   n (17) J. Cent. South Univ. (2016) 23: 544\u2212554 548 The upper and lower signs in Eqs. (16) and (17) correspond to generation of the concave and convex sides of the pinion tooth surface, respectively. The vector function and the unit normal of the pinion generating surface (b) p\u03a3 are similar to that of the gear generating surface (b) g\u03a3 (see Section 2.1.2), which are detailed no longer. 2.2.2 Equations of pinion tooth surface Coordinate system Sm{Xm, Ym, Zm} is rigidly connected to the cutting machine (Fig. 4). The top, bottom and middle of Fig. 4 are the machine front view, the machine bottom view and the side view (the projection of head-cutter). The cradle is rotated about the Ym-axis; the p1-axis is projection of the pinion axes in the XmOmYm-plane. The points P and O0 are the mean contact point and the center of the head cutter, respectively. O2 is the cross point of gear, and Om is the machine center. Some of the machine-tool settings of the given hypoid generator are: the machine root angle of pinion \u03b3m1, the machine center to back XP, the cradle angle of pinion q1, the radial distance of pinion Sr1, the blank offset Em1, the sliding base XB1, the swivel angle J1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure13-1.png",
        "caption": "Fig. 13. Four possible modes of relative micro-motion for a helical gear pair. (a) transverse translation; (b) misalignment of gear shaft due to rotation in the transverse plane; (c) mismatch of gear faces due to axial shift; (d) non-synchronous rotation between two gears.",
        "texts": [
            " For example, in a misaligned case represented by Fig. 12b, the penetrating condition only happens on the top end, therefore only the spring on the top is triggered and the force will be applied on two gears on the plane z = H/2 with opposite directions. For 3D gear dynamics in a CCHGP unit, there are four possible modes of micro-motions. The first mode is the transverse micromotion, for which the shafts of two gears remain parallel to each other, and the centers of the gear shafts shift in the transverse plane, as shown in Fig. 13a. The second mode is the misalignment that gives a non-parallel position of two gear shafts (Fig. 13b). The third mode is the axial micro-motion which gives the mismatch of two gear faces, due to the axial forces (Fig. 13c), and the fourth one is the non-synchronous rotation, or meshing error, between two gears (Fig. 13d). In this work, the focus is on the first three modes, and can be modeled by modeling the behavior of the journal bearings and the lateral lubrication gaps. In this work, it is assumed that Mz is fully balanced by the shaft torque, that driver and driven gears in the external gear pumps are rotating at constant angular speed without no transmission error, and the angular speed is unaffected by the loading. The study of the meshing error will be a focus of future work. Typical EGP designs are two-gear four-journal-bearing systems (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001468_iccas47443.2019.8971725-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001468_iccas47443.2019.8971725-Figure1-1.png",
        "caption": "Fig. 1. ATI system geometry.",
        "texts": [
            " Here, we introduce the statistical model of the multichannel ATI-SAR signal. Consider an ATI SAR system consisting of N antennas moving along direction x (azimuth) and separated along the azimuth by the two-way baselines bn from the first antenna, with n = 1, . . . , N and b1 = 0 [3]. Assume bn H , where H is the platform distance from the ground, and a target on the ground moving with a constant velocity vT = vxx+ vrr, where vx and vr are the azimuth and range velocity components, respectively (see Fig. 1). Let Z = [Z1, Z2, . . . , ZN ]T , ZC = [Zc1, Zc2, . . . , ZcN ]T , ZT = [ZT1, ZT2, . . . , ZTN ]T , and ZW = [ZW1, ZW2, . . . , ZWN ]T , the complex N -dimensional vectors representing the SAR processed images, the stationary ground clutter, the moving target, and the white thermal noise signals, computed in a fixed pixel, where T stands for the transpose. The target detection problem can be stated as a hypothesis testing problem, since it can be reduced to the choice between the two statistical hypothesis H0 and H1, corresponding to the absence and to the presence of a moving target, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003900_jsvi.1997.1323-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003900_jsvi.1997.1323-Figure4-1.png",
        "caption": "Figure 4. A multi-bearing-rotor system with rigid supports.",
        "texts": [
            " (5) The integration of equation (2) produces the oil film pressure distribution p. The principle of virtual work applied to the instantaneous oil pressure results in the hydrodynamic forces Hs acting on the flexible sleeve. The same principle applied to the chamber pressure pc yields the force vector Cs . The fluid in the chamber was assumed to be incompressible here. Modelling of the compressible control fluid in the chamber and its influence on system dynamics are presented and discussed in reference [15]. Figure 4 illustrates a multi-bearing rotor system with a rigid foundation. The system is statically indeterminate. Its configuration is defined by vector a. The rotor was treated as a free\u2013free body and was modelled by the FEM using Timoshenko beam elements. After matrix condensation using the Guyan method, the equations of motion for the transverse vibration of the rotor can be expressed in the form, Mr \u00b7 w\u0308 +Kr \u00b7 w=Hr +Qr +Fr , (6) where w is the absolute position vector of the rotor nodes, Mr and Kr are the mass and stiffness matrices of the rotor, Qr and Fr are static load vector and external excitation forces acting on the rotor, respectively, and Hr is the vector of hydrodynamic forces from the oil film"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000501_tla.2016.7437183-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000501_tla.2016.7437183-Figure5-1.png",
        "caption": "Fig. 5 muestra el isom\u00e9trico del veh\u00edculo el\u00e9ctrico con recuperaci\u00f3n de energ\u00eda donde 1) son dos vigas, 2) son dos vigas cortas, 3) es una placa delgada, 4) son ocho motores el\u00e9ctricos, 5) son ocho flechas, 6) son 6 ruedas, 7) son ocho generadores, 8) son ocho soportes de placa de los generadores, 9) son dos bater\u00edas. El veh\u00edculo el\u00e9ctrico tiene 8 ruedas independientes, cada generador el\u00e9ctrico se conecta mec\u00e1nicamente con la rueda usando un eje de transmisi\u00f3n, las ruedas y los generadores se conectan con la placa delgada. Las ruedas trabajan al mismo tiempo y van en la misma direcci\u00f3n. Celdas de capacitores se podr\u00edan usar para salvar la energ\u00eda el\u00e9ctrica de los generadores o se podr\u00eda usar para el alumbrado.",
        "texts": [],
        "surrounding_texts": [
            "Figura 5. Veh\u00edculo el\u00e9ctrico. Los estados son x1=im, x2=v, x3=\u03c9m, la entrada es u=Vem, y la salida es y=Vsg, el modelo din\u00e1mico es como sigue: x Ax Bu y Cx Du = + = +  (27) donde 1 2 3 x x x x    =      , 1 2 8 8 2 2 0 0 0 m m m m m m m m R K L L n nm m K B J J A F F r\u03c3 \u03c3  \u2212 \u2212   = \u2212   \u2212  , 1 0 0 mL B    =       , g m L LD c= \u2212 , ( ) ( )1 10g g m m L L m g g mL LC c R R K K c = \u2212 +  . La ecuaci\u00f3n (27) representa el modelo din\u00e1mico del veh\u00edculo el\u00e9ctrico con recuperaci\u00f3n de energ\u00eda. r es el radio de la rueda, v es la velocidad lineal, \u03c9 es la velocidad angular, Fn es la fuerza normal, m es la masa total de veh\u00edculo el\u00e9ctrico, \u03c32 es el coeficiente de amortiguamiento, im es la corriente de armadura del motor, Lm es la inductancia de armadura del motor, Rm es la resistencia de armadura del motor, Vam es el voltaje de armadura del motor, Vem es la fuerza contraelectromotriz del motor, K1m es la constante de fuerza contraelectromotriz, \u03c9m es la velocidad angular del motor, ig es la corriente de armadura del generador, Lg es la inductancia de armadura del generador, Rg es la resistencia de armadura del generador, Vsg es el voltaje de salida del generador, K1g es la contante de torque del generador, \u03c9g es la velocidad angular del generador, Jm es el momento de inercia del motor, Bm es el coeficiente de fricci\u00f3n viscosa, K2m=Kaux\u03c6m, K2m es el coeficiente de fuerza contra-electromotriz del motor. Se considera \u03c9m=\u03c9g debido a que el motor y el generador est\u00e1n conectados directamente. im=(1/c)ig donde c<1 es un par\u00e1metro seleccionado por el dise\u00f1ador. La Tabla III muestra los par\u00e1metros del veh\u00edculo el\u00e9ctrico con recuperaci\u00f3n de energ\u00eda. TABLA III. PAR\u00c1METROS DEL EJEMPLO 2. 2 2 2 1 6 2 3 2 1 Par\u00e1metro Valor Par\u00e1metro Valor 0.0225m 50 0.01s/m 0.050H 0.04kgm 98.1N 20kgm/A 24V 1.537V/RPM 0.06kg 5 10 H 1000 0.04kgm /rads 0.001 5 10 kgm /rads m m m m em m g g m g r R L J Fn K V K m L R B c K \u03c3 \u2212 \u2212 \u03a9 \u00d7 \u03a9 \u00d7 La representaci\u00f3n en espacio de estados del veh\u00edculo el\u00e9ctrico se reescribe como (6) donde dn=d es la perturbaci\u00f3n, un=u es la entrada de control, xn=[x1,x2,x3]T son los estados, y yn=y es la salida. Si se consideran los par\u00e1metros de la Tabla III, entonces las matrices del sistema (13) son: [ ] 3 7 1000.0 0 30.74 0 130.8 2.943 500.0 0 1.0 20.0 0 0 1.00000 0 5.0002 10 1.0 10 n T n n n A B C D \u2212 \u2212 \u2212 \u2212   = \u2212  \u2212   =  = \u2212 \u00d7  = \u2212 \u00d7 (28) La matriz de perturbaci\u00f3n es: [ ]1 0.050 0 0 T nE = (29)"
        ]
    },
    {
        "image_filename": "designv11_14_0002619_302-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002619_302-Figure6-1.png",
        "caption": "Fig. 6. Demagnetization curves of Alcomax I11 at - 15, -75 and - 180\" C, showing also the unit permeance lines for the dimension ratios used Saturation intensity 4rIJ, 14 100 G at + 15' C, 14 300 G at -75' C, 14 700 G at -180' C.",
        "texts": [
            " 0 c to room temperature after heating to -60\" C. On cooling there is a reversible increase in magnetization, the slope of the curve being always less than -0.03% per \"C but varying somewhat with magnets of different dimension ratio. The effect of alternating field stabilization is to eliminate the irreversible change on heating; the reversible changes are similar to those for fully-magnetized ellipsoids. D E M A G N E T I Z A T I O N C U R V E R E S U L T S The demagnetization curves of Alcomax I11 at 15, -75 and -180\" C are shown in Fig. 6. It will be seen that the curves cross at H values just over -500, close to the BH,,,,, point. The unit permeance lines of the five ellipsoids are also 93 -6C - 4 0 - 2 0 0 2 0 4 0 60 Temperature CC) Fig. 4. Alcomax 111. L/D 2.66 shown in this diagram. In the cases of Alnico and cobalt steel there is an increase in the coercivity as well as B-values as the temperature is (a) fully magnetized; (6) reduced 5 :d by alternating field. The effect of successive temperature cycles between room temperature and -65\" C is shown in Fig",
            " The effects found for Alcomax are also shown by Columax, i.e. Alcomax with a preferred crystal orientation. Tests taken on a Columax bar of dimension ratio 2/1 give curves similar to Alcomax 111 magnets having working points below the BN,,, point. There is a permanent loss of 2.3 % as a result of cooling to -70\" C and reheating to room temperature, but this loss is almost eliminated after a previous stabilization. lowered with no crossing of the curves. D I S C U S S I O N The demagnetization curves of Fig. 6 give the explanation of the magnetization temperature changes of Figs. 1 to 4. The intrinsic magnetization per domain increases with decreasing temperature, as indicated by the change of Bk and 4rI , in Fig. 6. Therefore, magnetized ellipsoids with working points on the upper part of the demagnetization curve increase in strength on cooling and decrease in strength on reheating, without change in domain orientation. For ellipsoids with working points on the lower part of the curve, cooling gives a similar increase in intrinsic magnetization Per domain, but, due to the reduced coercivity, the total magnetization at the working point is reduced. This change can only be achieved by partial domain re-orientation or boundan' 122 BRITISH JOURNAL OF APPLIED PHYSICS movement and such changes are mainly irreversible, giving a magnetization loss on reheating to room temperature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000464_j.isatra.2015.12.017-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000464_j.isatra.2015.12.017-Figure2-1.png",
        "caption": "Fig. 2. Defect on outer race.",
        "texts": [
            " Accounting for the fact that compression occurs only for positive values of \u03b4, \u03b3i which represents the contact state of ith rolling element \u03b4i is introduced as \u03b3i \u00bc 1 if\u03b4i40; 0 if\u03b4ir0; ( \u00f06\u00de The contact force is the sum of the contact force from each of the rolling elements and the total force along the x and y axes can be obtained as Fbx \u00bc XN i \u00bc 1 \u03b3iKb\u03b4i 1:5 cos\u03b8i; \u00f07a\u00de Fby \u00bc XN i \u00bc 1 \u03b3iKb\u03b4i 1:5 sin \u03b8i: \u00f07b\u00de The damping of rolling element bearing can be estimated using Kramer method [22] as cs \u00bc \u00f00:25 2:5\u00de 10 11Kl\u00f0N\u2219s=m\u00de; \u00f08\u00de where Kl is the linearized stiffness of the rolling element bearing, which can be obtained by Tamura method [23]. In the present analysis, the damping value of the bearing used is 200 N s/m. The most cause of bearing failures is the local defect, such as fatigue spall and pitting corrosion. The outer race defect can be modeled as a slight dent and the geometrical interpretation of the faults is shown in Fig. 2. The fault on the outer race is stationary, and the position of the dent on outer race is important. Analysis has been made for the Please cite this article as: Xu Y, et al. Active magnetic bearings used a ISA Transactions (2016), http://dx.doi.org/10.1016/j.isatra.2015.12.017 dent in the middle of the loading zone and at the bottom of the outer race, where outer race defect normally occurs [24,26]. The length of the dent is l and the following relationship can be established: d\u00bc rb ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi rb2 l 2 2 s ; \u00f09\u00de \u03c6\u00bc 2 sin 1 l \u00f02R\u00de ; \u00f010\u00de where rb is the ball radius; \u03c6 is the central angle of the outer defect; d is the max increment of the radial clearance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003720_s0039-9140(96)02040-1-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003720_s0039-9140(96)02040-1-Figure5-1.png",
        "caption": "Fig. 5. Microcell with inverted working electrode.",
        "texts": [
            " In Fig. 4, (1) is a glass capillary tube for the sample (10-40/tl), (2) is granular glassy carbon, (3) is pencil lead for contact, (4) is glass wool, (5) is the platinum auxiliary electrode, (6) is the reference electrode and (7) is the supporting electrolyte. The filling of the microcell with a sample was carried out by immersion of the tip in a sample solution. For emptying, the tip is touched to a tissue. This kind of solid working electrode, which is set from bottom to top, is illustrated in Fig. 5 [24]. The analogous state of the mercury working electrode is shown in Fig. 6 [42]. Working electrodes in the inverted state have been used in microcells more often than working electrodes in the usual state because of the possibility of applying a sample of smaller volume. Both macro [12,13,24,29,34,43] and micro [45 49,51,53] working electrodes in the inverted state have been used in microcells. The first development of the working electrode in the inverted state by Iwamoto et al. [24] was based on the use of a Pt plate (Fig. 5). On the surface of the Pt plate (1), the sample drop (50 /~1) (2) is placed. The Pt wire auxiliary electrode is in the form of a ring (3) and the tip (4) of the electrolytic bridge of the reference electrode is immersed in the sample drop. In the microcell (Fig. 5), a capillary with mercury as the working electrode that is directed upwards has also been used [24]. The cone form of the microcell with the mercury drop working electrode on the bottom of the cell is shown in Fig. 6 [42]. In Fig. 6, (1) is the Pt contact, (2) is the mercury drop of 5/~1 volume, (3) is the sample of 10/tl volume, (4) is the capillary of the electrolytic bridge of the auxiliary and simultaneously of the reference electrode and (5) is the capillary for the inlet of nitrogen. In the microcell (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001951_elmar49956.2020.9219027-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001951_elmar49956.2020.9219027-Figure1-1.png",
        "caption": "Figure 1. Kinematics of a quadcopter",
        "texts": [
            " Likewise in [14], leader\u2013follower type of fault-tolerant formation control approach for UAVs is proposed in the presence of potential collisions and actuator failures. This paper is organized as follows. The model preliminaries are described in Section II. The fault-tolerant control algorithm for the swarm of multi-agents is proposed in Section III. Simulation results are portrayed in Section IV, which is followed by the conclusion. In this work, the dynamics of each node in the swarm of multi-agent system is based on the nonlinear model of a quadcopter as shown in Fig. 1 [15]. The inertial reference is 978-1-7281-5973-7/20/$31.00 \u00a92020 European Union 62nd International Symposium ELMAR-2020, 14-15 September 2020, Zadar, Croatia 79 Authorized licensed use limited to: Carleton University. Downloaded on November 07,2020 at 12:08:13 UTC from IEEE Xplore. Restrictions apply. the earth shown as (x, y, z) that is the origin of the reference frame. The drone is assumed to be a rigid body that has the constant mass symmetrically distributed with respect to the planes (x, y), (y, z), and (x, z)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002378_0040517520936289-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002378_0040517520936289-Figure3-1.png",
        "caption": "Figure 3. Schematic illustration of the testing specimens in different tensile directions.",
        "texts": [
            " Consequently, tight weave areas undergo lesser shrinkage as compared to the loose weave areas, creating the differential shrinkage within the fabric structure. The higher shrinkage of the loose weave areas will force the tight weave areas to collapse and create double-directional folds in a parallel in-phase zig-zag fashion. To measure the NPR effect in five different tensile directions, the specimens of the auxetic woven fabric were subjected to a tensile test. The tensile directions included two principal directions and three biased directions. Figure 3 shows the schematic illustration of specimens in these tensile directions. Photos of test specimens showing the outlines of the unit cell in different tensile directions are presented in Figure 4. A square was marked on each specimen to facilitate the measurement of the size changes in both tensile and transverse directions during the tensile tests, as shown in Figure 5(a). The ASTM D5035-95 standard was followed and an Instron 5566 tensile testing machine was used to conduct a tensile test with the following parameters: gauge length (150mm); crosshead speed (50mm/min); pretension (0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003699_s0043-1648(98)00232-4-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003699_s0043-1648(98)00232-4-Figure3-1.png",
        "caption": "Fig. 3. Schemalie of seat sh~wing region intaged by the CCD camer:L A typical image captured by *,he CCD camera is shown with labels o.\" the ohservable salient features.",
        "texts": [
            " This average contact pressure was 8 MPa and the maximum sliding speed was 34.56 mm/s . The abrasive slurry in which the seal was operated was composed of a mixture of fireclay, bank sand, and water. To observe the wear process, a CCD camera was focused on the edge of the seal through the window. Images from the camera and information from a position encoder were transferred to a computer through a video capture board and digRally recorded. The 520 x 460 pixel images correspond to an area on the seal edge measuring 2.05 mm \u00d7 1.81 ram, Fig. 3. In the image, the abrasive slurry appears as a white substance near the top. 2.3. Laser induced florescence setup Laser-induced florescence (LIF) was used to visualize lubricant film thickness. This technique invoh, ed using a laser t0 fluoresce dye that has been dissolved in the oil. A low-pass filter is used to isolate the fluorescence of the dye from the background la,~er light, thus enabling the concentration of dye and hence oil to be measured [ 6 ]. The oil was dyed using Rhodamine 6-G dissolved in a dicMoromethanoi"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002572_j.apm.2020.12.020-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002572_j.apm.2020.12.020-Figure1-1.png",
        "caption": "Fig. 1. Virtual prototype of the parallel tracking mechanism.",
        "texts": [
            " In Section 5 , after the numerical simulation, dynamic response curves of some flexible generalized coordinates are analyzed. In Section 6 , the natural frequency sensitivity index of parallel mechanism is analyzed by direct differential method with the dimensional parameters and cross-section parameters of flexible beams as independent variables. Section 7 provides the summary of the whole paper. In the high-speed motion environment, the elastic deformation of parallel mechanism limbs will inevitably occur. Fig. 1 shows the virtual prototype of the 4-RSR&SS parallel tracking mechanism. The 2-rotational DoF 4-RSR&SS parallel mechanism is made up of a base, a moving platform, four identical RSR limbs and one SS limb. The four RSR limbs connect the moving platform and the base by rotational joints respectively and the SS limb connects centers of the base and the moving platform by spherical joints. S i ( i = 1 ~ 4) is the center of the spherical joint in the i th RSR limb. A i ( i = 1 ~ 4) is the center of the rotational joint in the i th RSR limb which is connected to the moving platform"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002816_j.optlastec.2021.107024-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002816_j.optlastec.2021.107024-Figure5-1.png",
        "caption": "Fig. 5. Sketch of the 3D benchmark with its supports.",
        "texts": [
            " The design concept of the chosen support structures is shown in Fig. 4. This linear geometry is extruded up to the part\u2019s surface. To ease their subsequent removal, all the supports are connected to the part by using teeth with height and width equal to 1.50 mm and 0.80 mm, respectively. All the dimensions are proportional to the distance dsl between the supports (Fig. 4). This design allows the effective removal of non-transformed powder before cutting all the supports. A sketch of the benchmark with its supports is displayed in Fig. 5. The distance dsl is one of the parameters of the experiment. We also investigate the inclination angle of the overhang (\u03b1oh), its thickness (toh) and its length (Loh). Table 1 summarises the considered levels for each factor. A full factorial DOE would require 135 tests [22]. In order to reduce the number of tests a D-Optimal reduction is used [23,24]. The DOptimality is reached by means of the Coordinate-Exchange algorithm [25]. A total amount of 20 tests and 4 repetitions are obtained. The Laser Beam Melting System SLM 250HL employs a diodepumped single mode ytterbium fibre laser, Yb:YAG YRL-400 (IPG Laser GmbH), with a maximum power output equal to 400 W"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003896_60.790917-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003896_60.790917-Figure1-1.png",
        "caption": "Fig. 1. Capacitors ac motors. (a) C-IM; (b) C-PMSM; (c) C-RSM",
        "texts": [
            " Manuscript submitted August 27, 1997; made available for printing February 18, 1998. In essence for the + - component and the C-IM, rotor flux orientation is used. Rotor axis orientation is used for C-PMSM and C-RM based on the idea that for constant rotor flux (steady-state) the IM works like a very good reluctancesynchronous machine [9]. Thus, the three ac motors become quite similar to each other. 11. BASIC CONFIGURATIONS AND COORDINATE TRANSFORMATIONS The three types of motors are represented in Figure 1, which shows that the two stator windings are displaced From each other by an angle 6. The above configurations are represented by orthogonal-axis models shown in Figure 2, where the auxiliary winding has been referred to the main winding. For generality we have included a capacitor in the main winding also. Using a power-invariant and mmf-invariant transformation, the following governing equations for the three motors are obtained. (1) -1 v - E, = Z,L] .*I (5) CO C = -ajk sin 6 + a cos 6 - k -2 d2 sin 6 0885-8969/99/$10"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002152_978-981-13-3549-5-Figure5.3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002152_978-981-13-3549-5-Figure5.3-1.png",
        "caption": "Fig. 5.3 Scientific instruments of the Curiosity rover and their locations [32]",
        "texts": [
            " There are a number of sensors and cameras on board which serve for autonomous navigation and perception of both the environment and rovers\u2019 status. They also carry a comprehensive suite of instruments dedicated to in situ observation and investigation with the help of robotic arms. For example, each of the Spirit and Opportunity rovers has five degrees-of-freedom arm called the instrument deployment device [30], four sets of stereoscopic cameras, three spectrometers, a Microscopic Imager and a Rock Abrasion Tool for cleaning and grinding rock surfaces [31]. Figure 5.3 shows the location of scientific instruments on the Curiosity rover [32]. The instruments carried can be divided into four categories: (1) the remote sensing instruments mounted on the mast\u2014Mast Camera (Mastcam) and Chemistry and Camera complex (ChemCam); (2) the scientific contact instruments at the end of the robotic arm\u2014alpha particle X-ray spectrometer (APXS) and Mars Hand Lens Imager (MAHLI); (3) the analytical laboratory instruments inside the rover body\u2014Chemistry and Mineralogy (CheMin) and Sample Analysis at Mars (SAM); and (4) the environmental instruments\u2014Radiation Assessment Detector (RAD), Dynamic Albedo of Neutrons (DAN), Rover EnvironmentalMonitoring Station (REMS) andMars Descent Imager (MARDI)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002765_s11071-021-06327-0-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002765_s11071-021-06327-0-Figure2-1.png",
        "caption": "Fig. 2 Actual engagement between a single-pin track and a sprocket (a) and three contact situations (b)",
        "texts": [
            " According to the geometry of the tooth groove profile, it is assumed to consist of three cambered surfaces: the left, bottom and right. These three surfaces are numbered 1, 2 and 3 in turn. Figure 1 shows a side view of a groove, which is bilaterally symmetrical. Oj is the center of the sprocket, rp the radius of its dedendum circle and ra the radius of its addendum circle. O1, O2 and O3 are the centers of the left, bottom and right surfaces, respectively. Their corresponding radii are r1, r2, r3. Obviously, r1 \u00bc r3. The actual engagement between a single-pin track and a sprocket is illustrated in Fig. 2a. Contact forces can be assumed to occur between track pins and tooth grooves. According to different contact positions, there are three contact situations: the left surface contact, the bottom surface contact and the right surface contact, as exhibited in Fig. 2b. Figure 2 illustrates that the cross section of the track pin is not a standard circle. It has an extra protrusion. In this investigation, the section of the track pin is axisymmetric, which is composed of two arcs and two straight lines, as shown in Fig. 3. The radius of the primary circle is r4, and the extent of arc P1P2 is 270 . The radius of the protuberant circle is r5. Straight lines P1P3 and P2P4 are the common tangent of the two circles. The detailed dimensions of TVMP are listed in Table 1. An appropriate contact model for cylindrical bodies is developed based on the Hertzian contact law, which is written as [28\u201330] FK \u00bc pE d 2 d 2 DR\u00fe d\u00f0 \u00de 1=2 d[ 0 0 d 0 8< : \u00f01\u00de where E represents the composite Young\u2019s modulus, 1 E \u00bc 1 v2a Ea \u00fe 1 v2b Eb (a = 1,2,3, b = 4,5) and the quantities Ea, Eb and va, vb are the Young\u2019s modulus and Poisson\u2019s ratio [31]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003542_pime_proc_1948_158_045_02-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003542_pime_proc_1948_158_045_02-Figure3-1.png",
        "caption": "Fig. 3. An Axle during Flange Contact",
        "texts": [
            " Whereas with new profiles there is two-point contact between wheel and rail while a flange is in action, with worn profiles, owing to the increased radius at the root of the flange, there is commonly single-point contact. The actual point of contact depends on the lateral flange force, and as speed increases, and the motion becomes more violent, the tendency is for the point to move farther down the flange; that is (rC--rd) increases with speed. Some useful conclusions can be drawn without either mathematics or experiment. Consider an axle rolling \u201ceast\u201d at a speed V, and let the flange of the \u201cnorth\u201d The Action of the Flanges. wheel be in contact with the rail (Fig. 3). Let the angle between the plane of the wheel and the rail be $, and let its value when the flange first hit the rail have been y, hereafter called the \u201cangle of incidence\u201d; the corresponding angle when the flange loses contact with the rail will be called the \u201cangle of reflection\u201d. If the angle of reflection at any flange impact is less than the angle of incidence, fiange action is tending to damp out the lateral oscillation. If, however, the angle of reflection is equal A RAILWAY AXLE 427 to, or greater than, the angle of incidence, flange action is sustaining, or even increasing the lateral oscillation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003964_elan.1140080709-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003964_elan.1140080709-Figure5-1.png",
        "caption": "Fig. 5 . Cyclic voltammograms obtained with HMDE in 02- and 1- saturated aqueous solutions containing 0.1 M NaOH and 0.5 M KCI in the absence (A) and the presence (B) of I mM Cyt.c(Fe3+). Potential scan rate: 500 mV sC'.",
        "texts": [
            " In the presence of SOD (10pM), the cathodic peak current became larger compared to that without SOD and the anodic peak current could not actually be observed except with high potential scan rates. This result is in accord with expectations based on Equation 3. The electrogeneration of 0; at the 1-adsorbed mercury electrode could be also confirmed by the reaction with ferric cytochrome c (Cyt.c(Fe3+)) in which 0, reduces Cyt.c(Fe3+) to its ferrous form [38-401 Cyt.c(Fe3+) + OT - Cyt.c(Fe2+) + O2 (4) In the presence of Cyt.c(Fe3+), we could, as expected, observe the increase in ip\" for the one-electron reduction of O2 to 0 7 i.e., the catalytic current (see Fig. 5) . The reaction of Equation 4 has been widely used for in vitro 0; assay [41]. These facts confirm previous results [4, 6,20, 221 and strongly demonstrate that 0, can be electrogenerated by the oneelectron reduction of O2 at the 1-adsorbed mercury electrode in aqueous media. In the following experiments to search for new surfactants for the electrogeneration of O;, we will examine by cyclic voltammetry,' the SOD-catalyzed disproportionation of OF which might be formed because of its unique specificity and high rate (it is almost diffusion-limited)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000447_jmes_jour_1960_002_027_02-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000447_jmes_jour_1960_002_027_02-Figure5-1.png",
        "caption": "Fig. 5. Tube ironing using four dividing planes",
        "texts": [
            " These equations are identical to those obtained by Hill (I) and Green (2) using the same configuration as a velocity J O U R N A L MECHANICAL E N G I N E E R I N G S C I E N C E Vol2 No 3 1960 at UNIV NEBRASKA LIBRARIES on June 5, 2016jms.sagepub.comDownloaded from UPPER-BOUND VALUES FOR THE LOADS ON A RIGID-PLASTIC BODY IN PLANE STRAIN 181 discontinuity pattern. The identity of the solutions in this particular case arises fiom the fact that the force-field method here reduces to 'block-sliding' over the dividing planes and there is, therefore, no essential difference between the two methods if friction is not present. A better upper bound for high reductions may be obtained from the pattern of dividing planes shown in Fig. 5a. There are now three variable points, b, c, and d, and minimization of p/2k is correspondingly difficult. For smooth die and plug, however, the analytical solution given above may be extended to this configuration and it can be shown that the best position of point c occurs when triangles abc and cde are geometrically similar. The reduction ratios from a to c and c to e are then each equal to . d ( h / H ) . Further minimization shows that the best upper wound is obtained when angles and #* are each equal to ilr, the corresponding optimum value for a total reduction J O U R N A L M E C H A N I C A L E N G I N E E R I N G S C I E N C E 2 ratio of z/(h/H)",
            " p A Cot2 # - B Cot $+C 2 K = D+cot $ where A = - I+- ;( :) B = (1-k) cot CL D = cot (a+tan-' p) The best upper bound is determined by differentiation: p/2k = 22/Ad(A.DZ+B.D+C)-2A.D+B REDUCTION ( I - i ) H a Comparison of upper- and lower-bound values. 1 Upper bound using four dividing planes. 2 Upper bound using eight dividing planes. 3 Upper bound for perfectly rough die, velocity discontinuity p = 0.05. p = 0.05. JOURNAL MECHANICAL E N G I N E E R I N G S C I E N C E As for tube ironing a better upper bound for high reductions can be obtained from a similar coniiguration to Fig. 5a. Making the same assumptions as in the case of tube ironing, an analytical solution is possible: p/2k = (l+m).p'/2k p'/2k is the best upper-bound value, obtained from the configuration of Fig. 7a, for a reduction ratio of 2/(h/H), and Cot (a+y)--COt e H cot (a+y)+cot * m=- jp y=tan- 'p Numerical results obtained from the above equations are shown in Fig. 8. In Fig. 8a, curves 1 and 2 are upper bounds determined using the configurations of Figs. 7a and 5a respectively. Curve 3 is obtained from a velocity discontinuity type of solution for a perfectly rough die, along which the material shears, and curve 4 is a lower bound calculated from the equation The effect of increasing coefficient of friction on the upperbound values obtained is shown in Fig",
            "\\/2 J(x+1)-2 cot a J O U R N A L M E C H A N I C A L E N G I N E E R I N G SCIENCE By differentiation, it may be shown that the optimum die angle a1 when shear takes place on the die surface is given by If the die angle is greater than al, a better upper bound is obtained fiom the configuration shown in Fig. 9c, for which the lowest upper bound is As for tube ironing and sheet drawing, better upper bounds for high reductions may be obtained from Upper bounds. 1 p = 0 . 2 p = 0.10. 3 Shear on die face. 4 With dead metal zone. Vo12 No 3 1960 at UNIV NEBRASKA LIBRARIES on June 5, 2016jms.sagepub.comDownloaded from configurations similar to that shown in Fig. 5a. If the best upper-bound value determined for an extrusion ratio of 1/(H/h) is p'/2k, then p/2k = (l+m).p'/2k For extrusion, m is the reciprocal of the corresponding factor for sheet drawing. If the material shears along the die surface or the configuration of Fig. 9c is used, then p/2k = 2.p'/2k. Numerical values obtained from these expressions are shown in Fig. 10. If the die angle is 45\" (Fig. lOa) sliding will normally take place over the die surface. The assumption that shear takes place does not lead to serious error, however"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000171_0954406219878741-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000171_0954406219878741-Figure6-1.png",
        "caption": "Figure 6. Axis system definition.",
        "texts": [
            " The model describes the motor response to an input command and produces a set of four motors\u2019 angular velocities. The propeller and tilt model in turn takes the motors\u2019 angular velocity to produce individual thrust and differential moments. The individual thrust vector is oriented according to the tilt angle inputs. Furthermore, this model includes the tilt mechanism dynamics. Finally, the total forces and moments are fed to a generic rigid body dynamic model which models the rotation and translation dynamics of the platform in 3D space. Kinematics. As shown in Figure 6, there are three axes: world axis Fw, a body axis indicated by F b frame of reference (FoR) and an axis system per each rotor indicated by F ri FoR. The body FoR is attached to the system body, moves and rotates with the body. The relationship between position time derivative in Fw (velocities in world FoR) and body velocities is expressed using the relation in equation (1) _PN _PE _PD 2 64 3 75 \u00bc Rbw u v w 2 64 3 75 \u00f01\u00de where Rbw is Euler Rotation matrix to transform the projection of a vector from F b to Fw"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001543_s40430-020-2234-5-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001543_s40430-020-2234-5-Figure2-1.png",
        "caption": "Fig. 2 The sketch of thermal expansion of tooth profile",
        "texts": [
            " (1)kT = 1 T = Fn \u0394te According to the above definition, the key to calculating the thermal stiffness of gears is to find the method to calculate the thermo-elastic coupling deformation of the meshing point. In this paper, the finite element method and analytical method were used. The flow chart of the calculation process is shown in Fig.\u00a01. 3 Analytical method of\u00a0thermal stiffness of\u00a0gears The calculation method of the thermal expansion of gear tooth profile has been provided by literature [25]. As shown in Fig.\u00a02, after the thermal deformation, point K on the tooth profile moves to the position of point K\u2032. In the coordinate system, rk\u2032 is the radius of point K\u2032, rk\u2032 = rk + \u0394rk. \u03c6k\u2032 is the angle between the radius rk\u2032 and y axis; \u03c6k\u2032 = \u03c6k + \u0394\u03c6k, where \u0394rk, \u03c6k and \u0394\u03c6k can be calculated using Eqs.\u00a0(2), (3) and (4), respectively. (2) \u0394rk = rb tn + r3 b (tb \u2212 tn) r2 b \u2212 r2 n \u2212 rb(tb \u2212 tn) 2(ln rb \u2212 ln rn) + \u0394tk (rk \u2212 rb) Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:150 1 3 Page 3 of 13 150 where \u03c6k is the angle between the radius rk and y axis, sk is the tooth thickness at point K, \u0394tk is the temperature rise at point K, which is the difference between the bulk temperature of the meshing point K and the initial temperature, tn is the temperature of the inner hole surface of the gear, tb is the temperature of the base circle, R is the radius of the pitch circle, rb is the radius of the base circle, rn is the radius of the inner hole, \u03b1k is the pressure angle at point K, \u03b1 is the pressure angle of the pitch circle and \u0394sk is the circumferential expansion at point K, \u0394sk = \u0394tk\u03bbsk/2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.33-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.33-1.png",
        "caption": "Fig. 10.33 The structure of the boiling barrel",
        "texts": [
            " (1) Boiling process and equipment Boiling process includes normal pressure and pressure boiling, of which pressure boiling can be classified into in-batch pressure and continuous pressure boiling. 1. Normal pressure boiling Normal pressure boiling is in-batch production. The drawbacks of this method include long production cycle and small production capacity. For example, the production cycle of 1#NC is up to 40\u201350 h. The loading amount per barrel is 1000\u2013 1500 kg NC (a small scale; 100\u2013120 kg/m3). The boiling barrel is made of stainless steel with a diameter of 3\u20134 m and height of approximately 2.15\u20132.75 m. The structure of the barrel is shown in Fig. 10.33. The boiling barrel consists of inlet (1), switch (2), exhaust port (3), material circulation pipe (4), steam heating pipe (5), wastewater discharge port (6), outlet (7), and insulation baffle (8). The exterior of the barrel is equipped with 60\u2013100 mm thick insulation layer. The bottom of the boiling barrel is a false bottom, where there are many staggering circular holes with a diameter of 5\u20136 mm for the filtration of wastewater in NC. To facilitate the circulation of boiling water in the boiling process, in the inside of the barrel four vertical tubes along the vertical interior wall of the boiling barrel are installed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000204_j.autcon.2019.102996-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000204_j.autcon.2019.102996-Figure15-1.png",
        "caption": "Fig. 15. Hob rotation adjustment: (a) angle measurements; (b) applied rotation to hub.",
        "texts": [
            " The calculated \u03b2 angles (within the interval of \u2212\u03b1/2 to \u03b1/2) are used to compute the correction angle for rotation of the hub. The \u03b2 angles can similarly be measured from element to notch; counterclockwise is considered positive. The smallest \u03b2 value (\u03b2S) is added to the largest \u03b2 value (\u03b2L) for one given hub, taking into consideration the positive and negative signs. The smallest value of \u03b2 may be the largest negative angle. The correction angle is the average between \u03b2S and \u03b2L. Then, the notch positions are all rotated in the direction of the correction angle. Fig. 15a shows an example of angle measurements with three elements (numbered 0, 1 and 2) and a hub with eight notches (\u03b1=45\u00b0). This figure shows the placement of the first notch vector aligned with the first line (line 0). The angles between that vector and all the elements in that node are measured about the plane normal to the hub axis. The angle for element 1 is 29.89\u00b0, and for element 2 it is 260.92\u00b0, measured from element to notch, following counterclockwise direction. From this measurement, only, it is possible to calculate all other angles required, since notch vectors are at a constant angular distance. The correction angle for this example is 7.57\u00b0 counterclockwise (Fig. 15b). The same procedure applies, automatically, to any other hub with a different number of notches. After placing the hub in each mesh node, we can generate the adjusted geometry of the interconnecting parts. The interconnecting part geometry subdivides into three subparts: subpart A of the geometry depends on notch location; subpart C depends on element location, and subpart B connects A to C. Fig. 16 shows these subparts. 3.3.8.1. Generation of subpart A. Subpart A depends on the geometry and positioning of the hub notch"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001542_lars-sbr-wre48964.2019.00069-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001542_lars-sbr-wre48964.2019.00069-Figure1-1.png",
        "caption": "Fig. 1. Horizontal schema of Leader-Follower approach (adapted from [11])",
        "texts": [
            " The reference robot is called the leader and the one following the leader is called the follower. Thus, there may be several leader and follower pairs, which allow for larger and more complex formations. The methodology applied in drones can be divided into horizontal control (latitude and longitude) and vertical control (altitude). The focus of the present work is in the horizontal control. In the horizontal control, we try to maintain the desired relative distance dd and the desired relative angle \u03b8d between leader and follower, as shown in Figure 1. To get this, the follower controller needs the leader\u2019s linear speed v and angular speed a, and the relative distance d and angle \u03b8 between the leader and the follower [11]. In this section, we present the control architecture developed for the formation methods proposed aiming to simplify their implementations and modifications. It consists of four layers, maintaining the formation algorithms separately. The first is the Control layer, which controls the drone motors aiming to reach the goal. The second layer is the Movement layer, for the definition of the goal position and goal orientation of the vehicle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001438_lra.2020.2969161-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001438_lra.2020.2969161-Figure6-1.png",
        "caption": "Fig. 6. Gripping-force measurement capacity of the forceps-driver. (a) Experimental setup. The gripping-force applied to the sliding-bars is measured by the F/T sensor. (b) The embedded sensor calibration result: the gripping-force as a function of the angular displacement of the rotator.",
        "texts": [
            " Although the proposed master device exhibits the low operational frequency characteristic, this characteristic may not be regarded as a serious drawback when considering the practical repetitive operation speed of the forceps in the robot-assisted surgery. In the experiments on the driving characterization and working performances evaluation of the forceps-driver, we focus on calibrating the force sensor module embedded in the device and assessing its gripping-force measurement capacity. For this experiment, we configure the experimental bench consisting of the F/T sensor and the two sliding-bars as shown in Fig. 6(a). When the forceps-handles are clamped by the rotation of the rotator and the forceps-tips are closed accordingly, both sides of the F/T sensor are compressed by the two sliding-bars on which the forceps-tips are anchored. In the experiment, the embedded force sensor module is calibrated based on the gripping-force data that is acquired by the F/T sensor and then plotted as a function of the angular displacement of the rotator (see Fig. 6(b)). The angular displacement data from the potentiometer is processed with first-order low-pass filter of 50 Hz cutoff frequency. Even though Authorized licensed use limited to: University of Canberra. Downloaded on April 29,2020 at 07:28:20 UTC from IEEE Xplore. Restrictions apply. the slight error is observed between the data sets, the graph shows that the embedded force sensor module can function to measure the gripping-force appropriately. Also, the result demonstrates that the gripping-force measurement capacity of the device is about 4 N"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001293_s00170-019-04519-y-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001293_s00170-019-04519-y-Figure3-1.png",
        "caption": "Fig. 3 Internal shear device",
        "texts": [
            " After that, the analysis was carried out using optical macro photography. The internal shear of thread was carried out by associating, with the tapped specimen, a steel screw of quality 8.8. The quality of the screw was chosen to ensure that its thread has a higher resistance than the specimen thread. The internal shear behavior was limited to the determination of the thread response to an axial displacement imposed on the screw-nut assembly. To hold the washer and carry out this test, a specific device similar to that used during thread development was utilized (Fig. 3). During the test, the tapped specimen was firmly held between two hollow parts having an inner bore coaxial to each other. Special attention was given to the centering of the tapped specimen with respect to the lower bore. To reduce bending of the cantilevered part of the tapped specimen during the test, the distance between the edge of the bore and the thread root has not exceeded 1 mm (Fig. 3). With these considerations, a proper characterization of the thread response despite the response of the entire specimen is globally achieved. After the free installation of the screw in the tapped specimen, the screw was animated with a progressive axial displacement without any other action until the destruction of the assembly. All tests were carried out on a tension-compression testing machine with a maximum load capacity of 50 kN. Each test was conducted until the extraction of the thread as a helical element"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure19-1.png",
        "caption": "Fig. 19. Features of Design of Scoop-Controlled Coupling",
        "texts": [
            " Large thrust-equalizing ports of the kind described are cut through the shell of the runner in special cases where the power and speed are high, or where the arrangements for locating the driving and driven shafts axially are such that it is considered desirable to use such ports. They have a decided influence on the characteristics of the coupling, and the standard practice is to use a runner without thrust-equalizing ports, and to locate the driving and driven shafts axially by means of a duplex ball thrust bearing with the flexible tie rod mounting shown in Fig. 19, p. 102. Scoop Tube Effects. The scoop tube coupling was constructed for many years with a scoop tube of admittedly crude form, since negligible advantage had been shown by early tests of improvements in its hydraulic design. More recently it was found advantageous to increase the rate of oil circulating through the cooler and thus reduce the water consumption, because of the higher heat transfer rate obtained. An improved scoop tube with thin walls and a taper tip was successfully introduced to deal with the required increase in the circulating pressure, and its capacity was proved by tests upon experimental couplings to be easily sufficient for the new requirements",
            "\" The shell of the runner is omitted, with the object of reducing the hydraulic thrust, and this offers advantages with the bearing arrangements shown in the drawing. On the other hand, during manufacture and erection of such a coupling it is rather easy for the runner to be dropped and for one or other of the radial projecting vanes to be cracked, with the risk of subsequent fracture when running under load. The vanes of the conventional design of runner are effectively protected by the shell and no problem is presented in dealing with the hydraulic thrust by such simple means as a duplex bearing with tie rod as shown in Fig. 19. The most remarkable feature of the design in Fig. 11 is that the scoop tube is * Soviet Boiler and Turbine Building, 1937, No. 7. 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 95 located inside the working circuit near the inner profile diameter, and carries a swivel head with a \u201cweather vane\u201d to ensure that it presents an open mouth to any liquid circulating in its neighbourhood. It appears that the coupling is designed to transmit 870 h",
            " When remote control is desired a small worm-geared servomotor is the most convenient means ; a spring-loaded clutch obviates the need for limit switches, and permits the scoop tube to be controlled directly by the hand lever if the servo-motor supply is off. Automatic control can be applied directly to the scoop tube lever, for example by a speed governor on the driven shaft, or by a torque motor energized from the main motor circuit if constant motor load is desired. If the pressure or some other function of the driven machine is to be governed, the necessary regulator can usually be applied directly to control the scoop tube, and heavy servo-mechanism is not necessary. In the standard design of scoop-controlled coupling, Fig. 19, the fol- 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 102 PROBLEMS OF FLUID COUPLINGS 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 103 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 104 PROBLEMS OF FLUID COUPLINGS lowing features have been incorporated in the light of experience with the earlier standard and experimental types already described. (1) The coupling is self-contained, having no external reservoir tank, and is completely independent of auxiliary power for the filling and emptying of its working circuit",
            " Quite astonishing cases of distortion of apparently rigid bedplates have been observed and ascribed to the settlement of foundations. Hence in the new design the driven machine is isolated by a flexible coupling, which, in conjunction with the angular flexibility of the fluid coupling, will render the driving motor and driven machine quite independent of possible alignment or external thrust troubles, such as can arise from bedplate distortion, differential expansion effects, unbalanced loads, or simply crude erection. A double-disk flexible coupling is shown in Fig. 19. In essence it comprises two steel disks united at their inner diameter, slotted radially to increase the flexibility, and each welded to a ring for convenient bolting to the flange of the driving and driven half-couplings. The torsional and lateral rigidity of such a coupling is satisfactorily high, whereas it is very flexible in raking up angular misalignment of the two shafts. The most obvious criticism of such a design is that fatigue of the double disks might be expected to occur, but it has been established by test that the amount of deformation and resultant stress in the resilient disks is so small, when accommodating the maximum degree of mis- 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000578_978-4-431-55879-8_3-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000578_978-4-431-55879-8_3-Figure3-1.png",
        "caption": "Fig. 3 a Robot i can help robot j to cover i (V i j )free. Thick solid and dashed lines represent the blocked and free components of Ai j respectively. b\u2013c Conditions i checks to find out if it has to help a common neighbor k to cover a portion of the region (V k j ) f",
        "texts": [
            " We say that two patches U and W are adjacent, if U \u2229 W contains a line segment in B jk (not necessarily A jk), for some j, k \u2208 IN ( j and k are not necessarily neighbors)2. The significance of two patchesU and W being adjacent is that a robot can move freely between these patches. The patches are created as robots explore the regions to be covered. We will discuss the process of constructing Si in steps. Scenario i. Patches in V i j , j \u2208 N (i) The robot i enters V i j \u2282 (Vj\\V j0 j )free, j \u2208 N (i), if and only if \u2203l \u2208 {1, 2, . . . , | A f i j |}, s.t. ri j (l) \u2229 Ab ji = \u2205. This condition is illustrated in Fig. 3a. This patch, say U1, is adjacent to V i0 i and is added to Si . Scenario ii. Patches in V k j , k, j \u2208 N (i), k \u2208 N ( j): If the robot i enters a patch U1 \u2286 V i j , it explores U1. If a portion U2 of V k j , k \u2208 N (i) \u2229 N ( j) is adjacent to U1, then robot i will find out if k can reach this portion of V k j . Otherwise, this portion of 2As U and W belong to free space, U \u2229 W is either \u2205 or a permeable line segment. V k j will be added Si . We will discuss the situations in which robot i should or should not cover a patch in V k j . Let Pi jk be the vertex common to Vi , Vj , and Vk . 1. Consider a scenario, as illustrated in Fig. 3c, where U1 \u2229 V k j is a single line seg- ment and Pi jk \u2208 U1. Let u jk(m) \u2208 Ab jk contains Pi jk (Such u jk(m) exists as Pi jk is assumed to be part of U2 adjacent to U1). 1a. If u jk(m) \u2229 A f k j = \u2205, as illustrated in Fig. 3c, a, then k can reach U2, and hence i will not cover it. 1b. Otherwise, as illustrated in Fig. 3c, b, k can not reachU2 and i should cover it. The robot i can check if Pi jk \u2208 U1 \u2229 U2, and ifU1 \u2229 U2 is a single connected piece, while physically exploring the boundary of U1. 2. Consider a scenario, U1 \u2229 V k j is a not a single line segment or Pi jk /\u2208 U1, as illustrated in Fig. 4. In such a scenario, robot i will not be able to decide if U2 needs to be added to Si or not only based on available information. The patch U2 is added to Si , only if, while physically exploring the boundary of U2, the robot i reaches a portion of A f k j "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003566_pime_proc_1945_153_010_02-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003566_pime_proc_1945_153_010_02-Figure7-1.png",
        "caption": "Fig. 7. Flange Thickness Gauge",
        "texts": [
            " Profiles of these tyres were taken after 8,000 miles and again after 21,000 miles, at which stage film records were made of the wheel motion, and it was noticed that the wear on the cylindrical tread tyre was distributed differently from that usually encountered on the standard 1 in 20 coned tyre. The cylindrical profile tended to wear higher up the flange than the coned profile, with the result that the wear was disproportionately great in the region tested by the standard flange-thickness gauge. This gauge (Fig. 7 ) which allows a minimum flange thickness of 4 inch at a radial distance of 6 inch from the top of the flange, would, if used on the experimental profiles, have condemned them when they still had a satisfactory contoup and were reasonably thick at the root of the flange. There arose in consequence the necessity of fixing, for these tyres, a permissible limit of wear which would achieve the maximum life consistent with reasonable limits of safety, but before this could be done, No. 2 tyre of the trailer compo was in fact reported by a carriage and wagon examiner as having a thin flange"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001460_s00170-019-04874-w-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001460_s00170-019-04874-w-Figure16-1.png",
        "caption": "Fig. 16 DPM concentration profiles of different schemes. a N2. b N3. c N7",
        "texts": [
            " The increase of chamber length L results in a decrease of the inclination angle, and the powder convergence chamber becomes \u201csmoother\u201d due to the smaller inclination angle; as the angle of reflection of the particles becomes larger, the probability of being rebounded during the movement will be lower. Therefore, the level L = 100 is considered the optimal scheme. From above analysis, the preliminary optimal structure parameter is inclination angle a (25\u00b0), exit width w (1.5 mm), chamber length L (100 mm). The powder concentration profiles of several schemes are selected to confirm the validity of the discussion and analysis results. Intuitively, as shown in Fig. 16, the concentration profiles of scheme N3 reveal better convergence within the target area and N2 is more uniform. This result is in good agreement with the above discussion. As the exit width w is the most significant factor impacting powder flow distribution, another group of design schemes are simulated to explore the detailed influence of the exit width on powder flow and obtain the final accurate optimal design. The number of parameter and their levels are listed in Table 5. Figure 17 presents the effect of the exit width w on the two indexes, Cxi and C\u03c3i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002721_j.aej.2021.01.012-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002721_j.aej.2021.01.012-Figure3-1.png",
        "caption": "Fig. 3 Globoid helical curve together with a coordinate system of the globoid worm gear.",
        "texts": [
            " The coordinates of the profile limiting points B andC can be obtained after assuming and determining the basic geometric parameters of the globoid worm gearing based on the AGMA or GOST standard and the equations presented in this paper [19]. In the case of the arc profile, the arc radius R must be assumed. The coordinates y0 and z0 have to be determined by using the parametric equation of the circle passing through two points, which in this case are points B andC. The equation parameters hp and hkcan be obtained based on trigonometric relationships as shown in Fig. 2. The parametric equation of the tooth flank surface of the worm hob is obtained by developing a tooth profile along the globoid helix (Fig. 3). The position vector of the tooth flank surface of worm hob with the rectilinear axial profile is determined by r 1 0\u00f0 \u00de 1 \u00bc M 1 0 1 M12 M2 0 2 M21 x1 u\u00f0 \u00de y1 u\u00f0 \u00de z1 u\u00f0 \u00de 1 2 6664 3 7775 \u00f04\u00de After development, one gets the following form: r 1 0\u00f0 \u00de 1 u1; u\u00f0 \u00de \u00bc x1\u00f0u\u00de cos u1\u00f0 \u00de a sin u1\u00f0 \u00de \u00fe a cos u2\u00f0 \u00de sin u1\u00f0 \u00de\u00fe \u00fey1\u00f0u\u00de cos u2\u00f0 \u00de sin u1\u00f0 \u00de z1\u00f0u\u00de sin u2\u00f0 \u00de sin u1\u00f0 \u00de x1\u00f0u\u00de sin u1\u00f0 \u00de a cos u1\u00f0 \u00de \u00fe a cos u1\u00f0 \u00de cos u2\u00f0 \u00de\u00fe \u00fey1\u00f0u\u00de cos u2\u00f0 \u00de cos u1\u00f0 \u00de z1\u00f0u\u00de sin u2\u00f0 \u00de cos u1\u00f0 \u00de a sin u2\u00f0 \u00de \u00fe y1\u00f0u\u00de sin u2\u00f0 \u00de \u00fe z1\u00f0u\u00de cos u2\u00f0 \u00de 1 2 6666666666666666666664 3 7777777777777777777775 \u00f05\u00de In the case of the tool with arc axial profile, the position vector of the surface assumes the following form: r 1 0\u00f0 \u00de 1 \u00bc M 1 0 1 M12 M2 0 2 M21 x1 h\u00f0 \u00de y1 h\u00f0 \u00de z1 h\u00f0 \u00de 1 2 6664 3 7775 \u00f06\u00de After development, one gets the following equation: r 1 0\u00f0 \u00de 1 u1; u\u00f0 \u00de \u00bc x1\u00f0h\u00de cos u1\u00f0 \u00de a sin u1\u00f0 \u00de \u00fe a cos u2\u00f0 \u00de sin u1\u00f0 \u00de\u00fe \u00fey1\u00f0h\u00de cos u2\u00f0 \u00de sin u1\u00f0 \u00de z1\u00f0h\u00de sin u2\u00f0 \u00de sin u1\u00f0 \u00de x1\u00f0h\u00de sin u1\u00f0 \u00de a cos u1\u00f0 \u00de \u00fe a cos u1\u00f0 \u00de cos u2\u00f0 \u00de\u00fe \u00fey1\u00f0h\u00de cos u2\u00f0 \u00de cos u1\u00f0 \u00de z1\u00f0h\u00de sin u2\u00f0 \u00de cos u1\u00f0 \u00de a sin u2\u00f0 \u00de \u00fe y1\u00f0h\u00de sin u2\u00f0 \u00de \u00fe z1\u00f0h\u00de cos u2\u00f0 \u00de 1 2 6666666666666664 3 7777777777777775 \u00f07\u00de The parameter u1 in the above equations denotes the worm hob thread length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001137_j.ijrmhm.2019.105069-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001137_j.ijrmhm.2019.105069-Figure12-1.png",
        "caption": "Fig. 12. Transverse rupture strength for different LBM parameter settings with 800 \u00b0C pre-heating.",
        "texts": [
            " The test results reveal that the network of cracks, occurring in the specimens without pre-heating, significantly corrupts the transverse rupture strength (T.R.S.). Due to the superposed impact of the cracks, a distinct correlation between the LBM parameters and the measured T.R.S. values is not identified for these specimens. The maximum measured values are slightly above 50MPa. In contrast, the crack-free samples generated with pre-heating show a significantly higher transverse rupture strength (Fig. 12). The maximum T.R.S. values as well as the lowest standard deviation are achieved for medium parameter settings. This can be attributed to the porosity (Section 3.1) as well as the microstructural features and phase compositions of the samples (Section 3.5). For low energy inputs, the porosity significantly increases due to insufficient wetting. For high energy inputs, the porosity increases due to entrapped process gases. With increasing energy inputs, undesired phases with inferior mechanical properties are additionally formed as a result of decarburization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003697_s0013-4686(98)00103-0-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003697_s0013-4686(98)00103-0-Figure1-1.png",
        "caption": "Fig. 1. Scheme of the \u00afow-cell (a) and the enzyme reactor placed in the lower part of the \u00afow-cell (b); (A) polypropylene pipet tip in which the carbon \u00aeber microelectrode is inserted; (B) block containing the Ag/AgCl reference electrode; (C) steel \u00afow outlet tube-auxiliary electrode.",
        "texts": [
            " A Linseis L 6512 recorder and a Metrohm E 510 pH-meter were also used. Cylindrical carbon \u00aeber microelectrodes (Union Carbide Corp., 8 mm o.d. and 4 mm length, s= 0.0010 cm2 ) were used as the electrode substrate to be coated with poly(3-methylthiophene). Batch measurements were carried out using a BAS MF 2063 Ag/AgCl, and a Pt wire directly immersed in the solution as reference and auxiliary electrodes, respectively. A 10-ml electrochemical cell from BAS, model VC-2 was also used. The scheme of the \u00afow-cell used is shown in Fig. 1. This con\u00aeguration consists of a modi\u00aecation on a T-shape cell design previously described in the literature [33]. The solution \u00afow passes perpendicularly through the microelectrode, which is inserted in a home-made methacrylate block provided with a 4-mm f \u00afow-channel (Fig. 1(a)). This device was designed in order to allow the whole active surface of the electrode to be in contact with the \u00afowing solution, and to assure the smallest possible dead volume. An easy replacement of the microelectrode was in this manner possible. When the enzyme reactor was used, it was placed in the lower part of the T-piece by means of adjustable plastic \u00aettings close to the microelectrode (Fig. 1(b)). Parts B and C of the scheme show the block containing the Ag/AgCl reference electrode, and the steel \u00afow outlet tube, which was also used as the auxiliary electrode. NADH, NAD+, L-lactic acid, 1-cyclohexyl-3-(2morpholinoethyl) carbodiimide metho-p-toluenesulfonate, and L-lactate dehydrogenase (LDH) from rabbit muscle (E.C. 1.1.1.27), 800 U mg\u00ff1, were obtained from Sigma. Other chemicals used were of analytical reagent grade, and water was obtained from a Millipore Milli Q puri\u00aecation system",
            " Various RVC cylinders (4-mm f, 5-mm length) were bored from a RVC block. Before use, RVC cylinders were activated as follows [35]: \u00aerst they were soaked in 6 M HCl with continuous stirring for one hour and washed repeatedly until neutral pH was reached. Then, RVC cylinders were immersed into anhydrous methanol with continuous stirring for two hours. Finally, they were oven-dried at 1108C overnight. Once activated, RVC cylinders were stored in a closed box until they were used. One activated RVC cylinder was inserted into the \u00afow cell as depicted in Fig. 1(b). Immobilization of LDH was accomplished in the following manner: 40 mg ml\u00ff1 carbodiimide solution in 0.05 mol l\u00ff1 acetic acid/acetate bu er solution of pH 5.1 was recirculated through the RVC cylinder at a \u00afow rate of 1 ml min\u00ff1 for 90 min. Then, the bu er solution, kept at 08C, was passed through the \u00afow-cell for 10 min. Next, a LDH (1 mg ml\u00ff1) in 0.1 mol l\u00ff1 H2PO4 \u00ff/HPO4 2\u00ff bu er of pH 7.0, also at 0o C, was recirculated for 6 h through the \u00afow cell at a rate of 0.2 ml min\u00ff1. Finally, the same phosphate bu er was passed through again for 10 min",
            " In both cases, negative relative errors higher than 10% were observed only for NAD+ or ascorbic-to-NADH ratios higher than 4:1, and errors lower than 5% were produced when these ratios were lower than 2:1. The fact of obtaining negative errors in the NADH steady-state current measurements, suggested that a slight fouling of the modi\u00aeed microelectrode surface occurred when high concentrations of NAD+ or ascorbic acid were involved. Flow-injection with amperometric detection Using the \u00afow-cell depicted in Fig. 1, a peak current versus applied potential plot was obtained from injections of 150 ml of a 1.0 10\u00ff5 mol l\u00ff1 NADH solution into the 0.05 mol l\u00ff1 H2PO4 \u00ff/HPO4 2\u00ff bu er of pH 7.0 used as the carrier solution. Although a similar shape to that of the steady-state current versus potential plot obtained by batch amperometry (see Fig. 4) was observed, a much better signal-tobackground ratio was achieved in this case. On the other hand, the in\u00afuence of an excess of NAD+ on the NADH response was also studied by this technique",
            "2) 105 nA mol\u00ff1 l when NADH was injected into the carrier containing NAD+, whereas the slope was (1.320.1) 105 nA mol\u00ff1 l when NAD+ was also present in the analytical solutions. As expected, the sensitivity decreased as the presence of NAD+ was more important. As an application of the CFME coated with P3MT as NADH sensor, a \u00afow injection method for the amperometric enzymatic determination of L-lactate has been developed. This method makes use of a RVC-based enzyme reactor inserted into the \u00afow cell (see Fig. 1(b)), which was constructed as described in the Experimental section. The performance of the \u00afow system containing the RVC reactor in the absence of the immobilized enzyme was tested by obtaining a calibration graph for NADH in the (1.0\u00b110.0) 10\u00ff5 mol l\u00ff1 concentration range. No signi\u00aecant di erences in the slope and intercept values were observed when comparing with the calibration plot obtained with the \u00afow system without the RVC reactor, the peak shape and the background noise being also similar in both cases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002799_tie.2021.3063991-Figure23-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002799_tie.2021.3063991-Figure23-1.png",
        "caption": "Fig. 23. Prototype of MH-PHEFM machine. (a) Stator. (b) Rotor.",
        "texts": [
            "14 Authorized licensed use limited to: Dedan Kimathi University of Technology. Downloaded on June 28,2021 at 18:41:49 UTC from IEEE Xplore. Restrictions apply. 0278-0046 (c) 2021 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. VI. EXPERIMENTAL VERIFICATION In order to validate the above theoretical analysis and simulation results, a prototype machine is manufactured and tested. Fig. 23 shows the main components of prototype machine. The inner stator with armature winding, FW and PM is illustrated in Fig. 23(a), and the rotor structure is shown in Fig. 23(b). The prototype machine is fabricated according to the parameters in Table III. Fig. 24 depicts the exploded view of the main parts of the proposed machine. It contains stator, windings, rotor, housing, bearings, output shaft, and fixed shaft. The main parts and assembly process are similar to the conventional PM machine. The tested and simulated open-circuit back-EMF waveforms at 2.5 A/mm 2 , 0 A/mm 2 , -2.5 A/mm 2 FW current are plotted in Fig. 25, where the rotation speed is 120 r/min. It is shown that the back-EMF waveforms are quite sinusoidal and the results match well"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001844_j.diamond.2020.108040-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001844_j.diamond.2020.108040-Figure1-1.png",
        "caption": "Fig. 1. a) Cubic sample geometry and orientation. The blue plane indicates the orientation of the prepared cross-sections and b) schematic illustration of the substrate heating concept. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " PBF-LB/M processing of the samples was performed using an M270 system by eos GmbH, Germany, equipped with a 200W Yb-fibre Laser (\u03bb=1064 nm). The utilized parameter sets are listed in Table 1. The nominal energy input per volume EV was quantified utilizing the volume energy density according to [16]: =E P h v dV L s s where PL is the laser power, h is the hatch distance, vs is the scan speed and ds is the layer thickness. The exposure was conducted in a rotated stripe pattern with a rotation angle of 67\u00b0 per layer. The cubic samples with an edge length of 10mm (see Fig. 1a) were processed under nitrogen shielding gas atmosphere. Microsections of the samples for light microscopy were prepared parallel to the building direction (z-axis) as indicated by the blue plane in Fig. 1. The influence of the substrate temperature was investigated at ambient temperature (296 K) and 473 K. For conductive heating, the substrate plate was mounted on four encased resistive heating cartridges (600W) (see Fig. 1b) and the temperature was monitored by two type K thermocouples. The porosity was measured according to the archimedean principle by immersing in ethanol with five repetitions per sample. A field emission electron microscope (FE-JSM 7001F, Jeol, Japan) employing energy dispersive X-ray spectroscopy EDS (Oxford Instruments, United Kingdom) was used to analyze the topography, morphology as well as the distribution of the particles and the elements. The HV 0.3 Vickers hardness of the diamond segments was determined according to DIN EN ISO 6507 using a hardness tester (Duramin-40, Struers)",
            " One can conclude, that the addition of diamond particles narrowed the parameter window in which densification can be assured and cracking is avoided. In this context, it is believed that the reduction of thermal gradients through a substrate heating and the decrease of laser power and scan speed is purposeful to reduce residual stresses and cracking, allowing a densification of up to 99%. After the fabrication of the diamond segments, XRD analyses were conducted on the metallographic prepared cross-sections of the diamond segments (in the YZ plane of the building direction, cf. Fig. 1). Due to the statistical distribution of the diamonds in the metallic matrix material and the incidental detection by means of XRD, it should be noted that the intensity of the reflexes of the Ni-coated diamonds was randomly distinct. Furthermore, when detecting diamond particles in the DMMCs, there were reflex superpositions with the matrix material at 2\u03b8= 44\u00b0, 75\u00b0, and 91\u00b0 (cf. Fig. 4). The diffractograms of the samples (Fig. 9), which were fabricated without substrate heating, exhibited the same preferred crystal orientation as the 316 L starting material (cf"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000844_1464419318819332-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000844_1464419318819332-Figure1-1.png",
        "caption": "Figure 1. RD interaction with IR.",
        "texts": [
            " The angular location of the individual roller by Shao et al.16 is represented as di di \u00bc 2 p 1\u00f0 \u00de N \u00fe c s\u00f0 \u00det, for IRD 2 p 1\u00f0 \u00de N \u00fe ct, forORD 8>< >: \u00f01\u00de where d1 is the circular arc length containing starting and ending boundaries of fault by Shao et al.16 d1 \u00bc 0:5d\u00f0 \u00de 2 0:5d Hd\u00f0 \u00de 2 rd d \u00bc J d \u00f02\u00de where d is the defect angle, d is the faulty race radius. For IRD, d \u00bc i. For ORD, d \u00bc o. Contribution of RD. For roller defect r \u00bc rt, r is the angular location of defect on roller with respect to reference. When RD strikes with inner race as denoted in Figure 1, the defect amplitude Bdri is specified as When RD strikes with OR as shown in Figure 2, the defect amplitude Bdro is specified as Bdro \u00bc r r cos J 2r sin d J mod r, 2 \u00f0 \u00de \u00f0 \u00de , 04mod r, 2 \u00f0 \u00de 5 d1 r r cos J 2r , d14mod r, 2 \u00f0 \u00de 5 d d1 r r cos J 2r sin d J mod r, 2 \u00f0 \u00de \u00f0 \u00de , d d14mod r, 2 \u00f0 \u00de 5 d 8>< >: \u00f04\u00de Bdri \u00bc r r cos J 2r : sin d J mod r, 2 \u00f0 \u00de\u00f0 \u00de , 04mod r, 2 \u00f0 \u00de5 d1 r r cos J 2r , d14mod r, 2 \u00f0 \u00de5 d d1 r r cos J 2r : sin d J mod r, 2 \u00f0 \u00de\u00f0 \u00de , d d14mod r, 2 \u00f0 \u00de5 d 8>< >: \u00f03\u00de where peak additional deflection Xr is described by Patel and Upadhyay18 for roller bearing as r \u00bc r r cos J 2r \u00f05\u00de Contribution of IRD"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002161_9783527813872-Figure3.7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002161_9783527813872-Figure3.7-1.png",
        "caption": "Figure 3.7 Schematic illustration for the fluoropolymer@BaTiO3 nanoparticles and the interfacial region within the nanocomposites. Source: Reproduced with permission of Yang et al. [208]. Copyright 2013, American Chemical Society.",
        "texts": [
            " A bimodal architecture containing a low density of long grafted polymer chains can provide sufficient entanglements with the polymer matrix, while the high density of short polymer brushes imposed a steric repulsion and screens particle/particle interactions [126, 134, 210]. That concept was introduced into polymer nanocomposites containing BaTiO3 nanoparticles and oligothiophene polymers [211], and the resulting nanodielectric polymer composites demonstrated superior \u201cbimodal\u201d particle dispersion and enhanced dielectric properties (Figure 3.7). However, in order to push the k further in the desired direction, a different inorganic filler should be chosen. Doped BaTiO3 can be a starting point [212]. PMMA-grafted (Ba0.94Ca0.06)(Zr0.16Ti0.84)O3 with inorganic loadings as low as 18 wt% displayed a high k\u223c 56 at 1 kHz and a low tan \ud835\udeff \u223c 0.02. The shape of the fillers is another physical aspect that can provide property improvement. Well-defined PS-grafted rutile nanorods allow a high percolative threshold up to 41 vol% before the formation of conductive networks (Figure 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002542_icem49940.2020.9270965-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002542_icem49940.2020.9270965-Figure5-1.png",
        "caption": "Fig. 5. Concentrate winding and flux distribution of open circuit in 12- slot/10-pole combination [10].",
        "texts": [
            " 2) Magnetic Isolation Without magnetic isolation, fault currents can induce large voltages in other healthy phases; meanwhile, the current in healthy phases can raise the fault current in the fault phase. Both the fault and healthy phases perform worse when mutual inductance exists. Magnetic isolation means the mutual inductance generated between windings is almost zero, and the designation of magnetic isolation can be achieved by slot/pole combination and winding configuration. The winding configuration of machine which can be divided into overlapping and non-overlapping. Adopting the fractionalslot concentrated-winding [10] (FSCW, Fig. 5), electrical machine can satisfy the requirement of magnetic isolation. Using the star of slots of designing single layer fractional1605 0slot concentrated winding has been investigated [11]. It is an effective method which based on the number of slot and poles to arrange the correct coil connection to maximize the main harmonics of EMF. Furthermore, properties of star of slots and the feasibility of magnetic isolation have been induced. Based on these conclusions, a dual three-phase 12- slot 10-pole PM motor has been analyzed and tested [12]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001019_rpj-07-2018-0182-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001019_rpj-07-2018-0182-Figure4-1.png",
        "caption": "Figure 4 Top and Bottom die set closed to clamp metal tube",
        "texts": [
            " Hydroforming a Metal bellow hydroforming Prithvirajan R. et al. Rapid Prototyping Journal D ow nl oa de d by N ot tin gh am T re nt U ni ve rs ity A t 0 2: 47 3 1 M ay 2 01 9 (P T ) single convolution involves four steps and they are repeated till the required number of convolution is formed. Step 1: The preformed metal tube is inserted over the rubber bladder and positioned at a required height where the convolution is to be formed (Figure 3). Step 2: Top and bottom die halves are closed laterally to clamp the metal tube with the rubber bladder (Figure 4). Step 3: Application of hydraulic pressure inside the bladder expands the metal tube (Figure 5). This step is also referred as bulging (Kang et al., 2007). Step 4: The top die is closed vertically downwards till it touches the bottom die, the metal tube gets the \u201cU\u201d shape convolution (Figure 6). This step is also referred as folding (Kang et al., 2007). Then the formed convolution is shifted down and all the four steps are repeated till the required number of convolutions are achieved. Design and process parameters adapted from the previous study (Prithvirajan et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000679_sysose.2016.7542934-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000679_sysose.2016.7542934-Figure4-1.png",
        "caption": "Fig. 4: Erle-Copter (Quadcopter)",
        "texts": [],
        "surrounding_texts": [
            "1) Private Cloud \u201dACE Fog\u201d: A ten server (Facebook\u2019s OpenCompute V1.0 spec), private cloud running Openstack Open Source Cloud Software is installed in the Autonomous Controls Engineering (ACE) Field Lab, which we have named ACE Fog. One of the ten servers provides control services to the cluster, which includes authentication, image creation and storage, management and networking services. The remaining nine servers provide a bank of storage and computational power where the virtual machines reside. Figure 1 shows a single server and a collection of servers in a wood rack. a) Configuration Specifications: The current configuration has the capacity for 216 VCPUs that can consume up to 493.5GB of RAM and 40.1TB of disk space. These figures account for overhead of the host server OS on the physical servers. The servers are all connected to a TrendNet TEGS16DG gigabit ethernet switch. b) Installation Method: Installation of OpenStack on the servers was accomplished via OpenStack Configurator for Academic Research (OSCAR) [9], which prepares OpenstackAnsible [10] for use on small clusters. 2) Network: In a system of robots requiring the processing of visual data, there needs to be a capable network fabric to support them. A Linksys AC3200 TriBand WiFi router was acquired for the wireless link to the mobile robots. This triband router provides 3.2Gbps total throughput on the wireless links. ACE Fog connects into the router via a wired connection with the TrendNet gigabit ethernet switch. Virtual compute nodes on ACE Fog can connect to the building network via software-based routing tools in OpenStack. 3) Software: ROS is used as an integrating tool in our testbed configuration. Communication between processes is handled via ROS messages. Distributed processing is currently being performed using ROS in a method similar in functionality to that of MPI. The multimaster fkie ROS package [11] is used to synchronize ROS topics across multiple machines. Namespaces are used to isolate individual robot\u2019s topics across all synchronized nodes in the system. For namespacing, a script on each robot computer sets the specific namespace via the ROS NAMESPACE system environment variable [12] before multimaster fkie and the appropriate driver nodes are started."
        ]
    },
    {
        "image_filename": "designv11_14_0002673_s40684-020-00283-7-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002673_s40684-020-00283-7-Figure7-1.png",
        "caption": "Fig. 7 Temperature distribution in the 135\u00a0W_1125\u00a0mm/s specimen before giving rise in the refill of melting flow",
        "texts": [
            " Of these 18 specimens, the 30\u00a0W_250\u00a0mm/s specimen is identified to be the sole one without keyhole in the melting zone. In the 2-D analyses for the temperature distributions in the SS316L powder bed and substrate, the curve with 1658\u00a0K is identified as the interface of the melting and solid zones in both the powder bed and steel substrate. The isothermal curve with 3150\u00a0K is identified as the interface of the keyhole and melting zone. An example of the temperature distribution arising in the 135W_1125 mm/s specimen is shown in Fig.\u00a07. The initial height determined by the refills of melting flows from all sides of the moving laser spot should be known before gathering the melting flows driven by the 1 3 surface tension forces to the keyhole in the solidification process. However, the initial height in the keyhole is unavailable in the present 2-D simulations. In this study, some refill heights with several microns higher than the powdersubstrate interface have been assumed to solve their final solidification profiles. The gatherings of melting flows in the solidification process for the 135\u00a0W_1125\u00a0mm/s specimen operating at t = 0, 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000937_s00419-019-01516-1-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000937_s00419-019-01516-1-Figure2-1.png",
        "caption": "Fig. 2 Schematic of the j th and kth contact spots a cluster of microcontacts",
        "texts": [
            " This observation is supported by the theorem [32] stating that if K is inverse-positive matrix (that is, K \u22121 exists and all entries of K \u22121 are nonnegative), then K is the so-called M-matrix and can be represented in the form K = k0 I \u2212 B, where k0 > 0, and B is a nonnegative matrix. Remark 3 (Effect of the micropad contact surface geometry on the microcontact pull-off force) The value of the pull-off force (12) obtained by Kendall [20] is valid for a frictionless flat-ended circular punch. It was found by Gao and Yao [18] that the pull-off force f \u2217 j is strongly influenced by the contact surface profile z = \u2212\u03d5 j ( \u221a (x \u2212 x j )2 + (y \u2212 y j )2/a j ), (16) where \u03d5 j (\u03c1) is the shape function of the j th micropad with the center at point (x j , y j ) (see Fig. 2), and \u03c1 is a dimensionless polar radius. In fact, following Yao and Gao [26], it can be shown that the JKR (Johnson\u2013 Kendall\u2013Roberts) prediction for the pull-off force at the j th microcontact is f \u2217 j = \u221a 8\u03c0 E\u2217a3 j \u03b3 \u2212 \u03c0 E\u2217 a j a j\u222b 0 r 1\u222b r/a j \u03c7 \u2032 j (t)dt \u221a t2 \u2212 r2/a2 j dr , (17) where according to the Galin\u2013Sneddon general solution [23,33,34], we have \u03c7 \u2032 j (t) = \u2212 2 \u03c0 d dt \u239b \u239dt t\u222b 0 \u03d5\u2032 j (x)dx\u221a t2 \u2212 x2 \u239e \u23a0 . (18) By substituting (18) into Eq. (17) and performing integration, we arrive at the formula f \u2217 j = \u221a 8\u03c0 E\u2217a3 j \u03b3 \u2212 2\u03c0 E\u2217a2 j 1\u222b 0 \u03d5\u2032 j (t)t 2dt\u221a 1 \u2212 t2 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001208_j.triboint.2019.105999-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001208_j.triboint.2019.105999-Figure3-1.png",
        "caption": "Fig. 3. Sealing ring with built-in skeleton and spring.",
        "texts": [
            " These sealing devices are simple in structure and have good sealing compensation properties, so they are widely used. \u2460 An elastomer is preset inside the sealing device. The elastomer and the sealing element form a combined seal, and the sealing force is provided by the compression deformation of the elastomer [1,2], as shown in Fig. 1. \u2461 The spring piece is put inside the sealing ring, and the resilience force generated by the deformation of the spring piece compensation for sealing contact pressure drop due to seal wear [3,4], as shown in Fig. 2. \u2462 In Fig. 3, the metal skeleton is embedded inside the sealing body, and a spring coil is arranged at a position opposite to the sealing lip [5]. The metal skeleton provides a large initial sealing force; the spring ring can compensates for the wear of the seal and keep the seal lip in good contact with the component. (2) Compensate for seal wear by auxiliary machinery device. The mechanical oblique-cone-slid-ring (OCSR) compensated seal designed by Hu et al. [6] The device can independently adjust the position of the axial bush, push the OCSR to squeeze the O-ring to deform and obtain greater contact stress, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000804_cgncc.2016.7828903-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000804_cgncc.2016.7828903-Figure1-1.png",
        "caption": "Figure 1, the vehicle is described by a right-handed inertial frame (E-frame) represented by x, y, z, and a right-handed body frame (B-frame) represented by bx , by , bz . Positive bx -axis points towards the front rotors (rotors 1 and 2), positive by -axis points towards left (rotor 1) and positive bz -axis is directed downwards. Positive sense of the three angular variables Roll ( ), Pitch ( ), Yaw ( ) is decided by a right handed rotation about positive x, y, and z axis respectively.",
        "texts": [
            " The mechanical structure of tricopter Tri-Rotor UAV mechanical structure includes a frame, three rotor arms, a steering gear and three sets of power system. And one of the rotors, tail rotor to be specific, is tilted by a suitable angle using a servo to get rid of the reaction torque problem[6]. Tri-Rotor has an advantage of having a rapid motion generated from its tilt rotor, which could also be the defect of this system since it requires a very accurate value of tilting angle for the stabilization of the system[7]. B. Establish the mathematical model for tricopter The simplified structure of Tri-Rotor UAV is shown in figure 1.We choose body frame when analyzing the force, and choose inertial frame when describing movement process. In 978-1-4673-8318-9/16/$31.00\u00a92016 IEEE The rotation matrix for Earth-frame g g g g gS O x y z\u2212 to Body-frame bS Oxyz\u2212 transformation can be expressed as ( , , ) cos cos sin cos sin sin cos sin cos cos sin sin cos sin sin sin sin cos cos sin sin cos cos sin sin cos sin cos cos R \u03c6 \u03b8 \u03d5 \u03b8 \u03d5 \u03b8 \u03d5 \u03c6 \u03d5 \u03c6 \u03b8 \u03d5 \u03c6 \u03d5 \u03c6 \u03b8 \u03d5 \u03b8 \u03d5 \u03c6 \u03d5 \u03c6 \u03b8 \u03d5 \u03c6 \u03d5 \u03c6 \u03b8 \u03b8 \u03c6 \u03b8 \u03c6 = \u2212 + + \u2212 \u2212 (1) In the process of actual flight, the flight characteristics of the Tri-Rotor UAV should be affected by many factors",
            " So we can make the following assumptions: \u2022 The elastic deformation of aircraft structure is small, and can be negligible. Treat the body as a rigid body, total mass is constant. \u2022 Product of inertia ,xy yzI I is approximately zero. \u2022 Aircraft center of gravity position can be adjusted in actual use, so as to the body coordinate system origin and we treat g as a constant value. \u2022 Flying at a low speed, the body of the aerodynamic drag is not big, and it can be ignored. In body frame, gravity G should be described as: 0 sin 0 cos sin cos cos B Earth body mg G S mg mg mg \u03c6\u03b8\u03c8 \u03b8 \u03b8 \u03c6 \u03b8 \u03c6 \u2212 = = (2) As shown in figure1, the tricopter uses the first arrangement of the three configurations. The three axis component for aircraft by the resultant force in body frame can be expressed as: 3 1 2 3 sin sin cos sin cos cos cos X mg Y F mg Z F F F mg \u03b8 \u03b4 \u03b8 \u03c6 \u03b4 \u03b8 \u03c6 = \u2212 = + = \u2212 \u2212 \u2212 + (3) The three-axis torque is expressed as 1 12 2 12 1 12 2 12 3 3 3 3 3 1 2 3 cos sin sin cos x x y y yz yz L F l F l M F l F l F l M N F l M M M \u03b4 \u03b4 \u03b4 \u03b4 = \u22c5 \u2212 \u22c5 = \u22c5 + \u22c5 \u2212 \u22c5 \u2212 = \u2212 \u22c5 \u2212 + + (4) Put the two systems of equations into the six degrees of freedom dynamic equations of aircraft[8-9], shown as below: 2 2 2 2 2 2 sin ( ) cos sin ( ) cos cos ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( E E E E E E E E E x yz zx xy y z y zx xy yz z x z xy yz X mg m u qw rv Y mg m v ru pw Z mg m w pv qu L I p I q r I r pq I q rp I I qr M I q I r p I p qr I r pq I I rp N I r I p q I q \u03b8 \u03b8 \u03c6 \u03b8 \u03c6 \u2212 = + \u2212 + = + \u2212 + = + \u2212 = \u2212 \u2212 \u2212 + \u2212 \u2212 \u2212 \u2212 = \u2212 \u2212 \u2212 + \u2212 \u2212 \u2212 \u2212 = \u2212 \u2212 \u2212 + ) ( ) ( )zx x yrp I p qr I I pq\u2212 \u2212 \u2212 \u2212 (5) Then we can see the relationship between angular velocity and angular acceleration in earth-frame, angular velocity in body-frame with X Y Z L M N"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure19.5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure19.5-1.png",
        "caption": "Fig. 19.5 Kinematic diagram of the exoskeleton: initial state of verticalization (a), intermediate state of verticalization (b). 1 lower leg, 2 thigh, 3 spine, 4 foot",
        "texts": [
            " As a result, at the output of the classifier of each channel, we obtain a number corresponding to the confidence in the command to rotate the exoskeleton servomotor. To aggregate decisions on the channels of classifiers, all outputs of the channel classifiers go to a fuzzy neural network; the defusifier of which generates a control signal to the servo motor controller. As a result of the analysis of this signal, the controller determines the speed and direction of rotation of the servomotor. A simplified kinematic diagram of an exoskeleton for a verticalization mode is shown in Fig. 19.5. For simplicity of calculations, we assume that \u03d51= 90\u00b0 and does not change during the verticalization process, and then the verticalization process can be carried out using two servomotors SM1 and SM2 controlling angles \u03d52 and \u03d53, respectively. We use five EMG channels that read the EMG signal from the gluteus maximus muscle, biceps femoris muscle, semi-membranous muscle, semi-tendon muscle, and major adductor muscle to control servomotors in the verticalization process (see Fig. 19.6). The verticalization process is controlled by decoding the EMG in the channels and the rate of verticalization depends on the intensity of the EMGsignals in the channels"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure29-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure29-1.png",
        "caption": "Fig. 29. Traction Coupling with Helix Ring Valve",
        "texts": [
            "UID COUPLINGS section for the active edge of the ring valve and the use of a large number of pressure-equalizing holes in the disk-shaped portion connected to the actuating push rod. Helix Ring Valve. In the larger units, auxiliary power is frequently used to operate the ring valve, but it is possible to make the coupling independent of such auxiliary energy by using the helix ring valve 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 123 design shown by Fig. 29. The ring valve is mounted on a quick-threaded nut working on a sleeve which surrounds the runner shaft and extends outside the casing, so that it can be held by a brake. When this is applied with the coupling running, its restraint rotates the sleeve through an angle of about 270 deg. relative to the shaft, thus moving the nut axially and closing the ring valve. A spiral torsion spring is connected between the sleeve and the runner shaft so that upon release of the brake the sleeve is rotated, and the ring valve returned to the open position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000127_s40034-019-00145-1-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000127_s40034-019-00145-1-Figure2-1.png",
        "caption": "Fig. 2 Positioning of gripper on nonwoven sample a in the center of the sample, b on the edge of the sample",
        "texts": [
            " The images of selected materials taken with Dino-Lite microscope under magnifications 509 and 2009 are presented in Table 2. For the purpose of grasping observations, nonwoven materials are cut into samples with dimensions 50 9 50 mm. After defining and preparing samples to be used in the investigation, the grasping of samples using vacuum grippers is commenced. In the procedure, each sample is placed on a flat surface (i.e., working desk) and a vacuum gripper is positioned. The two positions of gripper are defined: in the center of nonwoven sample and on the edge of the sample (Fig. 2). A pressure regulator is used to change the pressure entering into the ejector. The ejector inlet pressures are defined to 2, 3, 4 and 5 bar. After the material is grasped from the flat surface, it was raised to a height of 150 mm, as shown in Fig. 3. At this point, it was observed whether at a given input pressure the sample was raised to a defined height and whether afterward the material is still grasped by the gripper. The observations are noted in the table. The complete procedure is repeated for three different suction caps (flat, small circle and large circle), as well as for two types of ejectors (L and H)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002152_978-981-13-3549-5-Figure5.19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002152_978-981-13-3549-5-Figure5.19-1.png",
        "caption": "Fig. 5.19 Wheel\u2013soil contact mechanics [108]",
        "texts": [
            " The transformation matrix from { e} to { I} is Ae \u23a1 \u23a2\u23a3 Ct X1 \u2212At Bt \u2212(B2 t +C2 t ) tan \u03d5w X2 At X3 Ct tan \u03d5w X1 C2 t +A2 t +At Bt tan \u03d5w X2 Bt X3\u2212At \u2212Bt tan \u03d5w X1 At Ct tan \u03d5w\u2212Bt Ct X2 Ct X3 \u23a4 \u23a5\u23a6, (5.9) where X3 \u221a A2 t + B2 t + C2 t . Let eFe and eFe denote the wheel\u2013terrain interaction forces and moments, respectively, which act on the wheel in the coordinate { e}. The equivalent forces and moments that act on the wheel in the inertial coordinate { I} are { Fe Ae eFe Me Ae eMe . (5.10) (2) Wheel\u2013Terrain Interaction Mechanics: The soil applies three forces and three moments to each wheel, as shown in Fig. 5.19. The normal force, denoted by FN, can sustain the wheel. The cohesion and shearing of the terrain can produce a resistance moment MDR and a tractive force. A resistance force is generated because the wheel sinks into the soil. The combined net force of the tractive and resistance forces is called the drawbar pull FDP, which is the effective force of a driving wheel. When a wheel steers or the terrain is rough, the skid angle \u03b2 is generated to produce a lateral force FL, steering resistance moment MSR and overturning moment MO on the wheel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000360_978-981-32-9441-7-Figure19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000360_978-981-32-9441-7-Figure19-1.png",
        "caption": "Fig. 19. The photograph of the cracking cabinet from the early ground test",
        "texts": [
            " Metal rubber vibration isolator is selected as the vibration isolator. The impact test methods are according to the environmental test method of GJB150.18-86 for military equipment. Since the total weight of the anvil is less than 900 kg, the drop weight and the anvil stroke are executed see Table 7 below. Three pendulum impact tests were carried out in horizontal and inclined directions of 30\u00b0. During the grounding test before the structural optimization, cracking occurred on the rotating shaft of the top turntable of the cabinet (see Fig. 19). After the structural optimization, the pendulum impact test was successfully carried out (see Fig. 20). After the after-test investigation, the cabinet and internal equipment are intact and the electronic equipment works normally. Structural Performance Improvement and Optimal Design 53 The anti-shock environment of the ship-based equipment is harsh, and there are great risks in structural design. Structural optimization design based on simulated analysis has important referential significance for improving structural performance",
            " The 3D model of a common plugging-pulling set is shown in Fig. 17. The operating process is shown in Fig. 18, and the steps are listed as follows. a. Plugging. As shown in the direction marked with solid arrow, a ! b ! d ! e f. The solid lines in the figure show the force applied to the plugging-pulling set. b. Pulling. As shown in the direction marked with dotted arrow, f ! e ! c ! b a. The dotted lines in the figure show the force applied to the plugging-pulling set. a. Accuracy control on the radial dimension Figure 19 is a schematic diagram used to express the relation. To simplify the expression, the value of e and g is set as negative, and the f and h is set as positive. The coaxiality between the axis 1 and the axis 2 is i. Fig. 16. Schematic diagram of the location pin Fig. 17. 3D model of the plugging-pulling set Research on the Design Method of Blind-Mating for the Modules 149 The largest value of the radial deviation Dbetween the location 1 and location 2 used to install the connectors is D \u00bc b m 2 \u00fe i\u00fe d p 2 \u00femax\u00bdf g; h e The value D cannot exceed the value the connector could float",
            " Schematic diagram of microstrip antenna unit Electromechanical and Thermal Synthesis Analysis 225 As can be seen from the above figure, the resonant frequency of the antenna is 5.4 GHz, and the maximum echo loss is \u221241 dB. The antenna can match well at the central frequency. The radiation characteristics of the optimized radiation patch are good, and the maximum gain of the antenna patch is 9.32 dB. The array of antenna radiation elements studied in this paper is rectangular grid array, as shown in Fig. 19. The array antenna has a total ofM * N elements with a distance of 50 mm. Fig. 17. Echo loss curve of radiation patch unit a E surface b H surface Fig. 18. Radiation patch gain pattern 226 K. Cui et al. The methods used in this paper for calculating the electrical performance of array antennas are as follows, First, it is necessary to extract the displacement information of all nodes on the antenna array at the maximum moment of structural thermal deformation, and to fit the surface in MATLAB to generate the surface equation, then the deformation surface of microstrip array antenna after thermal deformation can be obtained",
            " Then the equivalent structure size of satellite-borne microstrip antenna element after thermal deformation is obtained by projection method, and the influence of thermal deformation on the electrical performance of antenna array can be analyzed by HFSS software. The fitted surface is shown in Fig. 20 below. The gain pattern of the E-plane and the H-plane of the spaceborne active phased array antenna with or without thermal deformation is shown in Fig. 21 below. Table 3 below is the data sheet for the influence of thermal deformation on the gain pattern of the microstrip array antenna. Observation point \u03b8 yd x d y z mnP \u03c6 N-1 0\u0302r O x 0 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5\u22c5\u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5 \u22c5\u22c5\u22c5 Fig. 19. Schematic arrangement of elements for rectangular grid planar array Fig. 20. Thermal deformation surface fitted by MATLAB Electromechanical and Thermal Synthesis Analysis 227 All in all, the thermal deformation of satellite microstrip array antenna structure reduces the electrical performance of the antenna. The main manifestations are as follows: gain decreases by 5.6336 dB; left first SLL decreases by 5.134 dB; right first SLL decreases by 5.3233 dB; beam pointing to left 0.005\u00b0. In this paper, the whole satellite system model of space-borne active phased array antenna is taken as the research object, and the electromechanical and thermal analysis is carried out",
            " In order to prevent the formation of tunnel hole, tack welding with low welding depth by FSW was applied, which diminish the tunnel holes effectively. Meanwhile, the incorrect design of cooling plate structure is another reason that induces the tunnel holes. When the lap bench was too narrow, the metal will flow into the flow channel which will induce the formation of the tunnel holes (Figs. 17 and 18). The flow channel under the welding line can also increase the risk of the tunnel holes. In Fig. 19, it can be seen that tunnel holes was formed since the channel under the welding line under the welding line make it difficult for the cooling plate to supply enough support during the FSW process. (a) proper width of the lap (b) narrow width of the lap Fig. 17. The cross section of the cooling plate with different width of the lap. Fig. 18. The cooling plate with flow channel under the welding line. 890 T. Wang et al. During FSW, the improper welding parameter will reduce the heat input and make the metal plasticized not enough, then the loosen defect will form"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003681_971510-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003681_971510-Figure2-1.png",
        "caption": "Figure 2. Finite element model of caliper",
        "texts": [
            " The rotor finite element model cons sts (1 4) of 1365 solid elements and 2678 nodes as shown in Figure 1 : It has been shown experimentally that the coefficient 01' friction p varies with many parameters such as contact pressure, sliding velocity and temperature. To model the friction forces accurately, the influence of these parameters on the coefficient of friction should be included. In our formulation, the coefficient of friction is expressed as: where / / y S / / is the rnagnitue of the sliding velocity. p is the contact pressure and 7 is the surface temperature. COMPUTATION PROCEDURE 'The finite element moldel of caliper consists of 2130 solid elements and 3294 nodes as shown in Figure 2: At each time step, the computation starts with solving the dynamic equations. of motion to obtained the displacements, velocities and accelerations at each node by using Equations 1 through 4. The variation of contact surfaces are then determined by checking and computing the penetrations The finite element model of pads and backing plates consist of 756 solid elements and 1182 nodes as shown in Figure 3: The finite element of anchor braket consists of 639 solid elements and 1224 nodes as shown in Figure 4: The brake system consists of 5390 elements and 9368 nodes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001357_j.oceaneng.2019.106812-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001357_j.oceaneng.2019.106812-Figure15-1.png",
        "caption": "Fig. 15. Comparison of cross section shapes after welding (deformed scale: 50).",
        "texts": [
            " 13 are all identical comparing with experiment. The inter pass temperature is considered as about 200 \ufffdC for later FE computation. Fig. 14 shows the contour plotting of highest temperature distribution as well as welded zone, it also can be seen that welding arc has a little bit shallower penetration in the thickness direction comparing to the thickness of cylinder. This kind of temperature distribution caused by welding will not shrink the examined cylinder perimeter but influence its shape. As show in Fig. 15, original cross section shape of examined cylinder with yellow color was compared to deformed shape caused by welding with orange color. It also can be seen that the cross section shape has welding distortion with shrinkage in upright direction and expansion in horizontal direction. From the above computational results and explanation, the mechanism of welding distortion generated during rack and cylinder joining on the cross sectional shape of cylindrical leg structure can be clearly understood"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000171_0954406219878741-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000171_0954406219878741-Figure3-1.png",
        "caption": "Figure 3. Single tilt-quadrotor platform photo.10",
        "texts": [
            "9 The feasibility of the platform with reasonable capabilities was discussed, and a list of requirements was driven for the avionics, and the platform is shown in Figure 2. The list includes the required measurements such as linear position and velocities (from visual tracking system), angular velocities (from gyroscopes), tilt angle and spinning velocity of rotors (motor controller and servo motor) and linear and angular accelerations. The presented hardware was discussed, which runs a real-time version of Ubuntu operating system. Another work group10 also developed a platform with different capabilities, as shown in Figure 3. The developed platform has a single tilt axis with 1.4 kg weight. The design was driven by the requirement to have a system capable of performing perpendicular hover. Therefore, the tilt mechanism was designed to have a wide angle, ranging from 0 to 260 . The platform has thrust to weight ratio of 1.75 and runs a microcomputer RX62T from Renesas Technology. A dual tilt-quadrotor platform11,12 was developed to have a platform with increased agility and reliability. The design was based on fusing three actuation mechanisms: gyroscopic torques, thrust vectoring and differential thrusting"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure23.3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure23.3-1.png",
        "caption": "Fig. 23.3 An example of the algorithm in: environment 1, V-REP (a); environments 2 and 3, ROS (b)",
        "texts": [
            " In this table, rewards can only be added to one element in each row,which allows to combine actions of any links. To evaluate the performance, the developed Q-learning algorithm was tested in various environments. To test the developed algorithm of deep Q-learning, three environments were developed that differ in the number of expected observations. In this study, for training deep neural networks, we used the Keras library and TensorFlow framework. In the first environment, the agent needs to build up trajectory between two points using the model of the PhantomX Pincher manipulator (see Fig. 23.3a). In this environment, discrete actions and continuous states were used. In this case, one initial point and one target point are considered. In this environment, ANN approximates the values of observation space; therefore, the accuracy of inverse kinematics solution decreases. External forces, such as gravity, act on manipulator links in addition to the forces, exerted by motors. When testing Q-learning algorithm in the first environment, memory space was limited to 50,000 states. 32 data packets were used for training, and discount factor was 0.9. The developed algorithm trained ANN model, input and hidden layers of which contained 32 neurons, and the output layer\u201412 (see Table 23.2). For twoother environments, a simulation based onRobotOperatingSystem (ROS) was developed (see Fig. 23.3b). In the second environment, control signals were sent directly to manipulator joints without taking into account external effects. This made it possible to increase the rate of environment reaction to actions and to increase inverse kinematics solution accuracy.Discrete actions and discrete stateswere used in this environment. This environment is designed to solve inverse kinematics problem in limited manipulator working range, which is represented by a plane. Moreover, several initial and one target points were considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001459_s42835-020-00350-8-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001459_s42835-020-00350-8-Figure6-1.png",
        "caption": "Fig. 6 Parameters of vector field histogram algorithm",
        "texts": [
            " Therefore, in this work, the vector field histogram algorithm is used to detect and avoid obstructing obstacles in the path of the quadrotor based on the sensors\u2019 readings. The polar density histograms to detect barrier location and proximity are computed depending on the range of the sensors\u2019 readings. The radius size of the quadrotor is R , the minimum turning radius of the quadrotor is S and safety distance is D are the initial parameters that should be fed to the VFHA for safely avoiding the obstacles blocking the desired path [39]. Figure\u00a06 shows all of those parameters. The cost function is computed to determine the final orientation of the vehicle. The following equation is used to compute the cost function: where a, bandc are constant, Td is the Target direction (alignment of vehicle path with the goal), Wo is the Wheel orientation (the difference between the next direction and current vehicle orientation), Pd is the previous direction (the difference between the previously selected direction and the new direction of the vehicle)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure32-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure32-1.png",
        "caption": "Fig. 32. Self-Aligning Ball Bearing with Resilient Mounting",
        "texts": [
            " Thus, any wear of the main engine bearings, or lack of accurate alignment, would cause heavy loadings and vibrations to be transferred to the outboard bearing and imposed on the internal roller and ball bearings supporting the runner shaft. If, as in many cases, the engine crankshaft is stiff enough to carry the coupling directly, it is usual to mount a flexible coupling between it and the driven machine. If an outboard bearing must be used with a 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 128 PROBLEMS OF FLUID COUPLINGS rigid type coupling, the inner race may be mounted upon a flexible rubber bushing of the \u201cSilentbloc\u201d type as shown by Fig. 32. This bearing should be packed up sufficiently to support the coupling and it gives a small degree of resilience to safeguard the internal bearings carrying the runner shaft from undue loading due to lateral vibrations, misalignment, or wear of the main bearings. 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 129 In other cases it is possible to adopt the \u201cflexible type\u201d of traction coupling which has no internal journal bearings and thus permits of both parallel and angular misalignment within the limits of the running clearances"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure4-1.png",
        "caption": "Fig. 4. Mx and My moments generated by three asymmetric patterns of fluid-pressure forces on the radial gear surface. (a) x-component; (b) ycomponent; (c) z-component because of the helix overlap of helical gears.",
        "texts": [
            " The force analysis includes the consideration for two types of forces experienced by the gear, coming from two sources: forces exerted by the pressure of fluids, and the contact force between driver and driven gears. Furthermore, the fluid-pressure forces experienced by the gear are divided into two different categories: the forces acting on the radial surface of the gear, and the forces acting on the axial surface of the gear. The X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 outcome of the force analysis gives three components of the force vector and the moment vector experienced by each gear: F\u2192= \u239b \u239d Fx Fy Fz \u239e \u23a0 M\u2192= \u239b \u239d Mx My Mz \u239e \u23a0 (5) Fig. 4 shows three different patterns of how these asymmetric loadings are generated by fluid-pressure forces acting on the radial surface of gears, which result in the Mx and My moments that are absent for spur gear pumps. For the force components on the transverse plane, due to non-uniform application area distribution along the axial direction, which leads to a distribution of forces, a moment with respect to x- and y-axis can be generated, (Fig. 4a and b); on the other hand, because of the helical gear structure, the fluids surrounding the gear tends to push the leading flank of one tooth to one axial direction and push the trailing flank of the adjacent tooth to the opposite direction (Fig. 4c), which creates a moment and a net axial loading. The axial rotor surfaces (Fig. 5a) in helical gear pumps are in direct contact with the axial lubrication gaps. The latter is the lubrication gap between the gears and the bearing blocks (Fig. 1e), typically with a thickness on the order of 10 \u03bcm. When the pump is operating and the gears are spinning at high speed, the thin layer of fluid film in these gaps provides the bearing function to support the load, as well as the sealing function to prevent the working fluids from leaking to the outside"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000578_978-4-431-55879-8_3-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000578_978-4-431-55879-8_3-Figure2-1.png",
        "caption": "Fig. 2 a The region bounded by dark lines is V i0 i . b Illustration of Ab i j and A f i j . c Illustration of V j i when Vi is repartitioned between robots j \u2208 N (i)",
        "texts": [
            " Distributed spatial partitioning problem: For each i \u2208 IN , the i th robot should construct Si , a contiguous collection of topologically connected cells, such that the collection S = {S1, S2, . . . , SN } partitions Qfree. Let V i0 i \u2282 Vi be the subset of Vi containing pi (0). If there are no obstacles within Vi , then V i0 i = Vi . The boundary of V i0 i is made up of portions of Ai j and obstacle boundaries. A point q \u2208 Ai j is reachable to robot i from pi (0), if q \u2208 V i0 i , and unreachable otherwise. Figure2a illustrates V i0 i with an example. Let Ab i j = {ui j (k)|Ai j \u2283 ui j (k) /\u2208 V i0 i }, where ui j (k)s are mutually disjoint con- vex sets, representing parts (line segments) of Ai j that are not reachable (blocked by obstacles) to the robot i . Similarly, let A f i j = {ri j (k)|Ai j \u2283 ri j (k) \u2208 V i0 i }, where ri j (k)s are mutually disjoint convex sets, representing parts (line segments) of Ai j that are reachable (not blocked by obstacles) to the robot i . See Fig. 2a for illustration. Note that A f i j = Ai j\\Ab i j . Let N f b(i) = { j |Ab i j = Ai j } \u2282 N (i).When j \u2208 N f b(i), entire Ai j is unreachable to the robot i ; then the robot i can not enterVj without enteringVk , for some k /\u2208 {i, j}. Let N b(i) = { j |Ab i j = \u2205} \u2286 N (i). Note that N f b(i) \u2282 N b(i) \u2286 N (i). Note that Ai j = Ab i j \u222a A f i j , and Ab i j \u2229 A f i j = \u2205, thus Ab i j and A f i j partition Ai j . If Ai j = Ab i j (that is, A f i j = \u2205), then we say that Ai j is impermeable to the robot i . If Ai j = A f i j , then we say that Ai j is fully permeable to the robot i . If Ab i j = \u2205 and A f i j = \u2205, then Ai j is partially permeable to the robot i . Note that Ai j = A ji , However Ab i j = Ab ji , and A f i j = A f ji , in general. Let V i j \u2282 Vj , for j \u2208 N (i), be a portion of Vj that would have been part of Vi with node set IN \\{ j}. See Fig. 2c for illustration. V i j = Vj \u2229 ( j V\u0303i ), where j V\u0303i is the Voronoi cell of i with nodes IN \\{ j}, or just N ( j). Each portion of Vi\\V i0 i , is part of V j i , for some j \u2208 N (i). If Ai j is fully impermeable to the robot i , that is, Ai j = Ab i j , then i will not be able to reach V i j . In this section, we explain the proposed distributed spatial repartitioning scheme. The i th robot first explores V i0 i and obtains the following information: (i) Vi , N (i), pi (0), pi (t), p j (0) \u2208 Q,\u2200 j \u2208 IN the position of itself and initial positions of all other robots; (ii) Ai j , Ab i j , and A f i j , for each j \u2208 N (i); (iii) the sets N f b(i) \u2282 N b(i) \u2282 N (i), and iv) V i0 i "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002091_0954405420978039-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002091_0954405420978039-Figure2-1.png",
        "caption": "Figure 2. Head-cutter with circular arc profile.",
        "texts": [
            " The modifications are introduced by the appropriate variation in machine tool settings and the geometry of the head-cutter. The machine tool settings are defined in Figure 1: sliding base setting (c), basic radial (e), basic machine center to back increment (f), basic offset Figure 1. Manufacture of pinion teeth. (g), tilt angle (b), and swivel angle (d). The other manufacture parameters are the velocity ratio in the kinematic scheme of the machine tool for the generation of the pinion tooth surface (ig1), the radius of the headcutter (rt1), and the radii of the tool profile (rprof1, rprof2, Figure 2). The pinion tooth surface is determined as it follows (Figure 1): ~r 1\u00f0 \u00de 1 =T4 T3 T2 T1 ~r T1\u00f0 \u00de T1 \u00f01a\u00de ~v T1, 1\u00f0 \u00de 0 ~e T1\u00f0 \u00de 0 =0 \u00f01b\u00de where ~r T1\u00f0 \u00de T1 is the position vector of head-cutter surface points, matrices T1, T2, T3, and T4 describe the coordinate transformations from system KT1 (attached to the head-cutter T1) to the coordinate system K1 (attached to the pinion). Equation (1b) describes mathematically the generation of pinion tooth surface by the head-cutter; ~v T1, 1\u00f0 \u00de 0 is the relative velocity vector of the head- cutter to the pinion and ~e T1\u00f0 \u00de 0 is the unit normal vector of the head-cutter surface. The equation of the head-cutter surface with straight line profile (Figure 1) is: ~r (T1) T1 (u, u)= u (rt1 + u tga1) cosu (rt1 + u tga1) sinu 1 2 664 3 775 \u00f02\u00de The equation of the head-cutter with circular arc profile (Figure 2) is: ~r T1\u00f0 \u00de T1 n, u\u00f0 \u00de= xTpi rprofi sinn rti cosu rti sinu 1 2 664 3 775 \u00f03\u00de where i=1, 2 is for the two circular arcs, and xTp1 = hTM + rprof1 sina1, yTp1 = rt1 + ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2prof1 x2Tp1 q , rt1 = yTp1 rprof1 cosn xTp2 = hTM rprof2 sina1, yTp2 = yTp1 rprof1 + rprof2 cosa1, rt2 = yTp2 + rprof2 cosn \u00f04\u00de Matrices T1, T2, T3, and T4 are defined by the following expressions: ~rT0 =T1 ~rT1 = cosb sinb 0 0 cosd sinb cosd cosb sind 0 sind sinb sind cosb cosd 0 0 0 0 1 2 664 3 775 ~rT1 \u00f05\u00de ~r0 =T2 ~rT0 = 1 0 0 0 0 sinc cosc e cosc 0 cosc sinc e sinc 0 0 0 1 2 664 3 775 ~rT0 \u00f06\u00de ~r01 =T3 ~r0 = cosg1 sing1 0 c cosg1 sing1 cosg1 0 f c sing1 0 0 1 g 0 0 0 1 2 664 3 775 ~r0 \u00f07\u00de ~r1 =T4 ~r01 = cosc1 0 sinc1 0 0 1 0 p sinc1 0 cosc1 0 0 0 0 1 2 664 3 775 ~r01 \u00f08\u00de where c1 =c10 + ig1 c c0\u00f0 \u00de For the velocity vector~v T1, 1\u00f0 \u00de 0 , it follows ~v (T1, 1) 0 =v(T) ig1 z (T1) 0 + g cosg1 z (T1) 0 ig1 z (T1) 0 + g sing1 ig1 y (T1) 0 sing1 x (T1) 0 c cosg1 h i y (T1) 0 2 6664 3 7775 \u00f09\u00de where ~r T1\u00f0 \u00de 0 =T2 T1 ~r T1\u00f0 \u00de T1 and ~e (T1) 0 =T2 T1 ~e(T1)T1 \u00f010\u00de For head-cutter with straight lined profile ~e(T1)T1 = sina1 cosa1 cosu cosa1 sinu 0 2 664 3 775 \u00f011\u00de For head-cutter with circular arc profile ~e(T1)T1 = sinn cosn cosu cosn sinu 0 2 664 3 775 \u00f012\u00de In the developed multi-objective optimization model the tooth contact pressure, the transmission error, and the energy losses in the gear mesh are minimized by appropriate variation of the machine tool settings and tool geometry"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000023_ipdps.2016.51-Figure19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000023_ipdps.2016.51-Figure19-1.png",
        "caption": "Figure 19. In a Mergeless Chain, good pairs do always exist.",
        "texts": [
            " 13, the run pairs are good pairs if the fat robots are both located on the same side. So, if no good pair existed, they must instead lie on alternating sides of the quasi lines. Fig. 18 shows how a sequence of horizontal quasi lines then must look like. Because we want to close the chain, we need a second vertical quasi line. If no merge is possible, this quasi line must point upwards, i.e., in an opposite direction than the first one. In this proof, we show by contradiction that if this always is the case, then the chain cannot be closed. In Fig. 19, we have grouped subchains of connected horizontal respectively vertical quasi lines and symbolize such a group by a dashed curve. For our argumentation, we take the robot s, which is the first robot of the first horizontal quasi line, as the starting point. We define an orientation vector which points from the second last robot of the last vertical quasi line to the second robot of the subsequent horizontal one. (If between these two quasi lines a stairway exists, then the orientation vector becomes longer than in the figure, but still points to the same direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000823_i2cacis.2016.7885310-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000823_i2cacis.2016.7885310-Figure1-1.png",
        "caption": "Fig. 1. Planar three-link kinematic chain",
        "texts": [
            " For a system S possessing n DOFs, there exist n dynamic equations of motion which can be written in the matrix form as \ud835\udc40?\u20d7\u0308? = ?\u20d7? + ?\u20d7? + \ud835\udc3a + ?\u20d7? , (6) where M is the mass matrix, ?\u20d7\u0308? is the acceleration matrix, ?\u20d7? is the vector of applied torques, ?\u20d7? is the vector of moments from centrifugal forces, \ud835\udc3a is the vector of moments from gravitational forces and is the vector of moments from external forces and torques. V. EXAMPLE This section presents the development of a dynamic model of a 2-dimensional three DOFs kinematic chain using Kane\u2019s method. The model is also verified by using Lagrange method later. Figure 1 shows the planar three link chain. In Fig. 1, N represents the ground reference frame, A the reference plane of link 1, B the reference plane of link 2 and C the reference frame of link 3. The system is fixed at joint A0 in N. \ud835\udf0f \ud835\udc41 \ud835\udc34\u2044 is the torque exerted by N on A, \ud835\udf0f \ud835\udc34 \ud835\udc35\u2044 is the torque exerted by A on B and \ud835\udf0f \ud835\udc35 \ud835\udc36\u2044 is the torque exerted by B on C. A. Method I: Kane\u2019s method The symbols in Fig. 1 represent the following = joints, = center of mass = segment A (link 1), = segment B (link 2) = segment C (link 3) where A*, B* and C* are the center of masses of segment A, B and C respectively.\ud835\udc5e1 is the counterclockwise rotation angle from \ud835\udc5b1\u0302 to \ud835\udc4e1\u0302, \ud835\udc5e2 is the counterclockwise rotation angle from \ud835\udc4e1\u0302 to \ud835\udc4f1\u0302 and \ud835\udc5e3 is the counterclockwise rotation angle from \ud835\udc4f1\u0302 to \ud835\udc501\u0302. \ud835\udc5b1\u0302, \ud835\udc5b2\u0302, \ud835\udc5b3\u0302, \ud835\udc4e1\u0302, \ud835\udc4e2\u0302, \ud835\udc4e3\u0302, \ud835\udc4f1\u0302, \ud835\udc4f2\u0302, \ud835\udc4f3\u0302, \ud835\udc501\u0302, \ud835\udc502\u0302 and \ud835\udc503\u0302are mutually orthogonal unit vectors. An anticlockwise circular arrow indicates a vector pointing out of the drawing and vice versa for the clockwise circular arrow. Hence, as depicted in Fig.1,\ud835\udc5b3\u0302, \ud835\udc4e3,\u0302 \ud835\udc4f3\u0302 and \ud835\udc503\u0302 are pointing out of the drawing. \u03c1A, \u03c1B and \u03c1C are distances of center of mass from their proximal ends. lA , lB and lC are lengths of the segments A , B and C respectively. Reference frames N, A, B and C are related by creating tables of direction cosines. Now, following the steps mentioned in Section IV: Step 1: To construct the direction cosines table e.g., for reference frame A with respect to the ground N (NRA), the following points should be taken into account: 1. A reference frame is affixed each body. During no motion, all basis vectors are aligned with the N basis vectors. When the segments move, the basis vectors move since they are rigidly fixed with the segment. 2. The plane of motion in Fig. 1 is defined by the basis vectors ?\u0302?1 and ?\u0302?2, with ?\u0302?2 pointing vertically upwards and ?\u0302?1 pointing in the direction of travel. As segment A moves through an angle of \ud835\udc5e1, as shown in the Fig. 1, the reference frame A gets aligned in the new orientation with respect to the reference frame N. A B C previous alignment basis vectors ?\u0302?1 and?\u0302?2. All the basis vectors have lengths equal to 1, the horizontal component of ?\u0302?1has magnitude cos \ud835\udc5e1 and points in the +?\u0302?1 direction. The vertical component has a magnitude of sin \ud835\udc5e1 and direction +?\u0302?2. Thus,?\u0302?1 = cos(\ud835\udc5e1) ?\u0302?1 + sin(\ud835\udc5e1) ?\u0302?2,?\u0302?2 = cos(\ud835\udc5e1) ?\u0302?2 + sin(\ud835\udc5e1) (\u2212?\u0302?1) and ?\u0302?3=?\u0302?3. Arranging these in the form of a table gives the table of direction cosines. Using the same steps, direction cosines tables of all the references frames relating them to each other are formulated in Table. II. The relationships between N, A, B and C basis (set of three noncoplanar vectors) vectors for the linkage in Fig. 1 is presented in Table II The elements of direction cosines table are the dot products between unit vectors of different coordinate reference frames. This facilitates the calculations done in Kane\u2019s method because of the regular need to compute dot products of vectors in different reference frames [4]. For instance, from Table. II, the dot product of \ud835\udc5b2\u0302 and \ud835\udc4e3\u0302 can be found just by finding the element in the \ud835\udc5b2\u0302 row and element in \ud835\udc4e3\u0302 column, that is, \ud835\udc5b2\u0302. \ud835\udc4e3\u0302 = 0. Step 2: Calculating angular velocities of reference frames, A, B and C with respect to the reference frame N, \ud835\udc41 \u0277 \u2192\ud835\udc34 = \ud835\udc5e1\u0307\ud835\udc4e3\u0302",
            " (41) These equations of motion can then be arranged in the matrix form of Eq. 6, and solving for T. B. Method II: Lagrangian method Lagrangian mechanics is based on the differentiation of the energy terms (kinetic and potential energies) with respect to the system variables and time. A Lagrangian can be defined as \ud835\udc3f = \ud835\udc3e \u2212 \ud835\udc43, (42) where L is the Lagrangian, K is the kinetic energy of the system and P is the potential energy of the system. Thus, for i=1\u20263, \ud835\udc47\ud835\udc56 = \ud835\udf15 \ud835\udf15\ud835\udc61 ( \ud835\udf15\ud835\udc3f \ud835\udf15?\u0307?\ud835\udc56 ) \u2212 \ud835\udf15\ud835\udc3f \ud835\udf15\ud835\udf03\ud835\udc56 , (43) Referring to the Fig. 1, the kinetic energy K of the linkage can be derived as \ud835\udc3e = ?\u0307?1 2 ( 1 2 \ud835\udc3c\ud835\udc34 \u2217 + 1 2 \ud835\udc3c\ud835\udc35 \u2217 + 1 2 \ud835\udc3c\ud835\udc36 \u2217 + 1 2 \ud835\udc5a\ud835\udc34\ud835\udf0c\ud835\udc34 2 + 1 2 \ud835\udc5a\ud835\udc35\ud835\udc59\ud835\udc34 2 + 1 2 \ud835\udc5a\ud835\udc35\ud835\udf0c\ud835\udc35 2 + 1 2 \ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc34 2 + 1 2 \ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc35 2 + 1 2 \ud835\udc5a\ud835\udc36\ud835\udf0c\ud835\udc36 2 +\ud835\udc5a\ud835\udc35\ud835\udc59\ud835\udc34\ud835\udf0c\ud835\udc35\ud835\udc362 +\ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc34\ud835\udc59\ud835\udc35\ud835\udc362 + \ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc35\ud835\udf0c\ud835\udc36\ud835\udc363 +\ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc34\ud835\udf0c\ud835\udc36\ud835\udc3623) + ?\u0307?2 2 ( 1 2 \ud835\udc3c\ud835\udc35 \u2217 + 1 2 \ud835\udc3c\ud835\udc36 \u2217 + 1 2 \ud835\udc5a\ud835\udc35\ud835\udf0c\ud835\udc35 2 + 1 2 \ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc35 2 + 1 2 \ud835\udc5a\ud835\udc36\ud835\udf0c\ud835\udc36 2 +\ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc35\ud835\udf0c\ud835\udc36\ud835\udc363) + ?\u0307?3 2 ( 1 2 \ud835\udc3c\ud835\udc36 \u2217 + 1 2 \ud835\udc5a\ud835\udc36\ud835\udf0c\ud835\udc36 2) + \ud835\udc5e1\u0307\ud835\udc5e2\u0307(\ud835\udc3c\ud835\udc35 \u2217 + \ud835\udc3c\ud835\udc36 \u2217 +\ud835\udc5a\ud835\udc35\ud835\udf0c\ud835\udc35 2 +\ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc35 2 +\ud835\udc5a\ud835\udc36\ud835\udf0c\ud835\udc36 2 +\ud835\udc5a\ud835\udc35\ud835\udc59\ud835\udc34\ud835\udf0c\ud835\udc35\ud835\udc362 + \ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc34\ud835\udc59\ud835\udc35\ud835\udc362 + 2\ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc35\ud835\udf0c\ud835\udc36\ud835\udc363 +\ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc34\ud835\udf0c\ud835\udc36\ud835\udc3623) + \ud835\udc5e2\u0307\ud835\udc5e3\u0307(\ud835\udc3c\ud835\udc36 \u2217 +\ud835\udc5a\ud835\udc36\ud835\udf0c\ud835\udc36 2 + \ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc35\ud835\udf0c\ud835\udc36\ud835\udc363) + \ud835\udc5e1\u0307\ud835\udc5e3\u0307(\ud835\udc3c\ud835\udc36 \u2217 +\ud835\udc5a\ud835\udc36\ud835\udf0c\ud835\udc36 2 +\ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc35\ud835\udf0c\ud835\udc36\ud835\udc363 +\ud835\udc5a\ud835\udc36\ud835\udc59\ud835\udc34\ud835\udf0c\ud835\udc36\ud835\udc3623), (44) and the potential energy P as \ud835\udc43 = \ud835\udc5a\ud835\udc34\ud835\udc54\ud835\udf0c\ud835\udc34\ud835\udc601 +\ud835\udc5a\ud835\udc35\ud835\udc54(\ud835\udc59\ud835\udc34\ud835\udc601 + \ud835\udf0c\ud835\udc35\ud835\udc6012) + \ud835\udc5a\ud835\udc36\ud835\udc54(\ud835\udc59\ud835\udc34\ud835\udc601 + \ud835\udc59\ud835\udc35\ud835\udc6012 + \ud835\udf0c\ud835\udc36\ud835\udc60123)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001938_icra40945.2020.9196853-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001938_icra40945.2020.9196853-Figure2-1.png",
        "caption": "Fig. 2. Proposed dry friction model. Selected friction direction maximizes energy dissipation under anisotropic friction constraints.",
        "texts": [
            ", \u2212fv), subject to other constraints (such as friction coefficients). We thus introduce a new smooth model for anisotropic dry friction that obeys the Maximum Dissipation Principal, formulated as: fenv dry =\u2212mg [ \u00b5l 0 0 \u00b5t ][ sin(arctan( \u00b5t vt \u00b5lvl )) cos(arctan( \u00b5t vt \u00b5lvl )) ] (8) Eq. (8) is derived by maximizing the dissipation \u2212fv with the assumption that the anisotropic friction is represented by a friction ellipse with the major axis equal to \u00b5t and the minor axis equal to \u00b5l as shown in Fig. 2. B. Viscous friction For underwater undulatory locomotion, anisotropic viscous friction plays an important role in generating forward motion. At low Reynolds number (i.e. small scale organisms in water), it can be considered the dominant reaction force. A model of reaction forces that comprises viscous friction exclusively can be considered as an approximation for small scale swimmers, such as tadpoles. We note that viscous friction has occasionally been used in the literature to model overland snakes as well [2], even though it is a less accurate approximation of dry Coulomb friction than the box model discussed above"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000506_s11771-016-3101-5-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000506_s11771-016-3101-5-Figure3-1.png",
        "caption": "Fig. 3 Circular profile blade and generating revolution surfaces for pinion head-cutter: (a) Illustration of the circular profile blad; (b) Generating tool surface for concave side; (c) Generating tool surface for convex side",
        "texts": [
            " The duplex helical method is based on the application of a helical motion of the cradle, on which the head cutter is mounted, with respect to the gear blank. Rotation of the cradle is accompanied by an infeed motion of the sliding base, on which the work spindle is mounted. 2.2.1 Generating surface for pinion circular profile head- cutter The geometry of the circular profile blade for the pinion head-cutter is represented in this section. Figures 3(a), (b) and (c) show illustration of the circular profile blade for the gear head-cutter, the generating tool surfaces for the concave and convex sides, respectively. As shown in Fig. 3, the coordinate system S0{X0, Y0, Z0} is rigidly connected to the pinion head-cutter, Z0 is the axis of the head-cutter, \u03b81 is the rotation angle, each side of the blade generates two sub-surfaces denoted as parts (a) and (b) of the generating surfaces, the segment of the circular arc (part (a)) with the curvature radius R1 (including the curvature radius of the concave side R1d and the convex side R1c) generates the working part of the gear tooth surface, the circular arc of radius \u03c11 generates the fillet of the pinion tooth surface (part (b))"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002903_j.engfailanal.2021.105453-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002903_j.engfailanal.2021.105453-Figure5-1.png",
        "caption": "Fig. 5. Longitudinal profile of the sample 3\u20131.",
        "texts": [
            " Engineering Failure Analysis 126 (2021) 105453 failed domestic sample, but it did not fail. To find out the reasons, we used sample 2# as the control group. Fig. 3(c) shows that the rubber near the axle is cracked and Fig. 3(d) shows that the surface of the cracked rubber is covered with sticky material. In order to determine the depth of the crack, two cracked samples were taken for the longitudinal cutting (denoted as 3\u20131) and ring cutting (denoted as 3\u20132). The cutting method is shown in Fig. 4 and the results are shown in Fig. 5 and Fig. 6. As displayed in Fig. 5, there are two symmetrical cracks on the upper and lower surfaces of the rubber joint, and the cracks are deep. By observing Fig. 6, it could be found that the cracks existed in all of the four rings, indicating that cracks had penetrated the upper and the lower sides. The test results of chemical composition of the axle and the steel cover materials of samples 2# and 3# are listed in Table 2 and Table 3. According to ASTM standard[19], it can be confirmed that the axle of sample 2# is 4140 alloy steel, the axle of sample 3# is 1045 carbon steel, and the steel covers of samples 2# and 3# are all 1020 carbon steel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001914_j.matpr.2020.08.486-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001914_j.matpr.2020.08.486-Figure1-1.png",
        "caption": "Fig. 1. The recast measurement samples and tensile geometries processed by WEDM.",
        "texts": [
            " Afterwards, the specimens are etched with Schanz prior to the investigation through optical microscope so as to reveal the effects of varying parameters on the thickness recast layer. The recast measurements is conducted based on the continuous thickness identified on the surface of each sample at multiple locations and average value is calculated. Similarly, the final geometry of dog bone shaped tensile samples are extracted from 125 26 3 mm prism based on the varying processing parameters and all surfaces are machined by Wire EDM, as seen both in Fig. 1. Based on the different number of passes, the experimental design and process parameters are chosen for sample processing in a single pass for Inconel 939. In this way, different energy discharges are achieved in a single pass with the parameters used. The surface morphology of specimens was examined through Polytec TMS-500 profilometer by extracting the three dimensional topography with a cut-off length of 0.8 mm at multiple locations. The area used for 3D surface analysis on gauge length of tensile has dimensional area of 22 mm 6 mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002222_j.surfcoat.2020.125371-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002222_j.surfcoat.2020.125371-Figure15-1.png",
        "caption": "Fig. 15. Nucleation and growth of TiC, (b) formation of W2C, (c) formation of the isolated W, and (d) formation of the regular cellular-shaped TiC layer.",
        "texts": [
            " demonstrated that the WCP reacted with the liquid Ti and started the 2WC+2Ti\u2192W2C+2TiC reaction because it had the maximum negative free energy [13]. The surface of the WCP provided numerous preferential heterogeneous nucleation sites for the TiC in this reaction. It should be noted that the cooling rate of the liquid around the WCP varied due to many factors, such as the particle size, position, and initial temperature of the WCP, which greatly affected the morphology characteristics of the WCP/Ti interfacial reaction layer. Gan et al. indicate that the highest cooling rate was at the top of the laser molten pool [18]. As displayed in Fig. 15(a), under the condition of a high cooling rate in the top part of the laser molten pool, the nucleation rate of the TiC was high as well, and many TiC grains were formed and encircled the WCP. The formation of TiC extracted C atoms from the surface of the WCP and transformed the WC to W2C. Furthermore, as Fig. 15(c) shows, in the gap between two adjacent TiC grains, liquid Ti contacted the W2C layer, and the W2C+Ti\u2192 2W+TiC reaction occurred when the temperature was below 2364 K. In Fig. 15(d), the TiC reaction products filled the gaps between two adjacent TiC, and a compact regular cellular-shaped TiC layer formed around the WCP. The W reaction product was trapped by the W2C/TiC interface, and isolated W was formed in the regular cellular-shaped reaction layer. The formation of a compact TiC layer around the WCP prevented the liquid Ti from making direct contact with the W2C layer and prevented further degradation of it. In Fig. 16(a), the TiC nucleation rate was low due to the low cooling rate conditions in the middle and bottom of the laser molten pool"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000222_0954406219885979-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000222_0954406219885979-Figure1-1.png",
        "caption": "Figure 1. The schematic of the double-row self-aligning ball bearing.",
        "texts": [
            " Then the formulation of the 3 3 stiffness matrix is analytically derived for the bearing. A comparison between the bearing stiffness of the proposed model and that from published literature is conducted to validate the proposed model. Finally, the effect of angular misalignment on the stiffness of the self-aligning ball bearing is studied systematically. Modeling of the double-row self-aligning ball bearing The double-row self-aligning ball bearing consists of inner ring, outer ring, cage and two rows of balls, as shown in Figure 1. The outer ring is a portion of a sphere, so the shaft can tilt around x-axis and y-axis freely. In general, there exists an internal clearance between balls and raceways. The internal clearance plays a very important role in the bearing performance, hence it should be considered in the bearing modeling. The following assumptions are made to establish an analytical model for the double-row self-aligning ball bearing: 1. The inner ring, the outer ring, and the balls are rigid except for the contact zones; 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002671_j.jsv.2021.115967-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002671_j.jsv.2021.115967-Figure1-1.png",
        "caption": "Fig. 1. An accelerometer is a single node with six-degrees-of-freedom and it can measure translational acceleration in one, two, or three principal directions (also known as uni-, bi-, and triaxial).",
        "texts": [
            " The acceleration of the rigid body is given by y\u0308 (t) = [ y\u0308 a (t) \u03b8\u0308(t) ]T (1) where y (t) is the full displacement vector, t is time, y a (t) is the translation displacement corresponding to the location of accelerometers, and \u03b8(t) is the angular displacement of the sensors. The accelerometers measures the acceleration, s (t) , as a summation of the correct acceleration, y a (t) , and a tilt error, \u03b5 (t) . s (t) = y\u0308 a (t) + \u03b5 (t) (2) For simplicity, we will derive the following equations for a triaxial accelerometer, s (t) \u2208 R 3 , but the derivation is expansible to any network of accelerometers. Thus, we have a triaxial accelerometer mounted in a single point, measuring the translational acceleration in the main directions, see Fig. 1 . In the case of no rotation of the triaxial accelerometer, we express the tilt error as the static tilt error. \u03b5 (t) = g , for \u03b8(t) = 0 (3) where g = [ 0 0 \u2212g ]T is a vector accounting for the gravitational acceleration (sometimes called earth\u2019s gravitational field vector). Next, we will introduce rotation around the x -, y -, and z-axes - denoted as \u03b8x , \u03b8y , and \u03b8z , see Fig. 2 , and we organise these rotations in a vector. \u03b8(t) = [ \u03b8x (t) \u03b8y (t) \u03b8z (t) ]T (4) We express the error as the accelerometer tilts by multiplying with the rotation matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002409_j.jfranklin.2020.07.028-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002409_j.jfranklin.2020.07.028-Figure1-1.png",
        "caption": "Fig. 1. Definition of tracking errors on the path.",
        "texts": [
            " For a matrix A , its complex conjugate transpose is denoted y A \u2217, H e (A ) =: A + A \u2217, He (A ) =: A + A \u2217 2 , tr [ A ] denotes the trace of matrix A , and A \u22a5 denotes he orthogonal matrix of A. For a real symmetric matrix, A > 0( \u22650) and A < 0( \u22640) denote ositive(semi) definiteness and negative (semi) definiteness. L 2 means the Hilbert space of quare integrable functions with the following norm: \u2016 v(t ) \u2016 2 = { \u222b + \u221e 0 v \u2217(t ) v(t ) dt } 1 2 . . Vehicle modeling and problem formulation In this paper, a two-degree-of-freedom (DOF) lateral dynamics model of four-wheel steerng (4WS) vehicles is considered [34,35] , see Fig. 1 , and the state-space form of the lateral ynamics is described by \u02d9 (t ) = Ax(t ) + Bu(t ) + B d d(t ) (1) here x(t ) = [ e y (t ) , e \u03c8 (t ) , \u03b2(t ) , r(t )] T is state vector with e y ( t ) being the distance between he center of the gravity of the vehicle and the reference path, and e \u03c8 (t ) = \u03c8(t ) \u2212 \u03c8 des (t ) is he orientation tracking error, \u03b2( t ) is the sideslip angle, r ( t ) is the vehicle yaw rate, u(t ) = \u03b4 s the front-wheel steering angle. d(t ) = [ d 1 ( t ), d 2 ( t ), d 3 ( t ), d 4 ( t )] T is disturbance representing the modeling error and external disturbance, the other system matrices are A = \u23a1 \u23a2 \u23a2 \u23a3 0 v x v x 0 0 0 0 1 0 0 a 11 a 12 0 0 a 21 a 22 \u23a4 \u23a5 \u23a5 \u23a6 , B = \u23a1 \u23a2 \u23a2 \u23a3 0 0 b 1 b 2 \u23a4 \u23a5 \u23a5 \u23a6 B d = diag { 1 , v x , 1 , 1 I z } , b 1 = C f mv x , b 2 = C f l f I z a 11 = \u2212C f + C r mv x , a 12 = \u2212 ( 1 + C f l f \u2212 C r l r m v 2 x ) a 21 = C r l r \u2212 C f l f I z , a 22 = \u2212C f l 2 f + C r l 2 r I z v x where m is the total mass of vehicle, I z is the yaw moment of inertia, C f and C r are the cornering stiffness of front and rear tires, l f and l r are distances from the center of the gravity to front tires and rear tires"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003005_j.procir.2021.05.142-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003005_j.procir.2021.05.142-Figure6-1.png",
        "caption": "Figure 6: Design of the Sensor Integrating Gear",
        "texts": [
            " Tests with white noise over a frequency band of 1 to 1000 Hz showed that there is no inaccuracy area in terms of frequency or amplitude distortion. Furthermore, there is no significant difference in phase and sensitivity. The MEMS accelerometer was capable of measuring small amplitude changes and showed linear response with an accuracy of \u00b1 3% in this case. Design: Two of the MEMS acceleration sensors, named MEMS1 and MEMS2, were mounted on the front face of one of the spur gears with sensing axes in opposite polarity using double-sided adhesive tape (Figure 6). For tangential acceleration measurements of the pinion, both sensors were oriented with sensitive axes (\u201cY\u201d) perpendicular to the axis of rotation (Figure 6). Position accuracy was ensured using an alignment gauge, fabricated by means of laser cutting and placed relative to the shaft of rotation. As stated before, the MEMS sensors are connected to a microcontroller by wires. The microcontroller and battery are mounted in a specifically designed housing which is mounted on the drive shaft (Figure 6). The housing is 3D-printed and the unbalance is reduced by design to reduce vibrations. The measurement strategy intends for 5 minutes of measuring and 55 minutes of sleep mode per hour which preserves energy. The changes in the gear vibration signal due to wear are slow enough in our test scenario that measurements every hour are sufficient to track the changes. With this strategy the battery lasts > 1000 hours of operation until recharging is necessary. Signal post processing: For a pseudo tach signal [13] the third harmonic of the gear mesh was used to determine a pulse signal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure39-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure39-1.png",
        "caption": "Fig. 39. Avonmouth Lock Gate Operating Winch",
        "texts": [
            "comDownloaded from and because of the flat shape of the characteristic curve it is essential that the overall efficiency of the worm reduction gear, including the losses in the vertical shaft bearings of the capstan head, should be closely estimated before the design of the coupling is settled, The pioneer capstan application was carried out in 1937 at the Avonmouth Docks, where it became necessary to replace the hydraulic operating machinery for the lock gates. Great care was taken to estimate the efficiency of the gear train, so that the stalled torque of the coupling 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COI'PLINGS 139 would be exactly correct when converted into the maximum desired pull in tons at the chain cable. A photograph of the installation appears in Fig. 39, Plate 4. ( 5 ) Conveyers. Another drhe where stalling is sometimes involved is that of assembly line conveyers. These machines are frequently required to be driven from three or more points, and the fluid coupling is used to equalize the load between the driving motors and to prevent heavy overload of the gearing and conveyer links, in the event of an obstruction jamming the conveyer line. Motors of 23-5 h.p. are commonly used, and being provided with torque-limiting fluid couplings, the whole construction of the conveyer can be light, as compared with the use of a single driving motor, which necessitates a heavy construction throughout to withstand the high pull required to drag the loaded conveyer around the many corners"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003126_s42835-021-00852-z-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003126_s42835-021-00852-z-Figure7-1.png",
        "caption": "Fig. 7 Thermal analysis of the one pole pair oil-cooling air-core stator: a is CFD model of the stator; b is temperature distribution contour of the stator; c is temperature rise distribution along Z direction at one slot region of a stator coil",
        "texts": [
            "K , respectively, whereas the axial and radial conductive coefficients of the back iron are 1 and 35 W\u2215m.K , respectively. (3) Computational fluid dynamics (CFD) analysis of the one pole pair air-core stator A successful mechanical design of the air-core stator is to bear both mechanical strain and thermal strain under operation conditions and to maintain the coils within acceptable levels of strain. Therefore, CFD analysis is essential before the thermal-stress analysis. One pole pair is taken as the solution domain based on the periodic of the HTS generator\u2019s structure and load, as shown in Fig.\u00a07. The two-phase heat flow coupling calculation of solid and cooling-oil is carried out in the CFD analysis. The coil with cooper conduct and insulation in the slot region is built as an entity with an anisotropic equivalent thermal conductivity as previously calculated. It is noted that the equivalent thermal conductivity of the coil in the end region is set in cylindrical coordinates, and the other parts are set in cartesian coordinates. And the K\u2212 turbulence model is selected because the flow transition in the model is smooth and there is no strong swirl and flow around",
            " The thermal expansion coefficient of the oil is input and the fluid and energy equations are opened until the continuous equation converges. The convective heat transfer coefficients in the slot region and the end region are about 100 and 60 W / m2 .K by the CFD analysis, respectively. Oil velocity in the end region is lower than that of the slot region because of the bigger space volume at the ends of the stator. The temperature distribution contour and temperature rise distribution along the Z direction at one slot region of the stator are shown in Fig.\u00a07. The maximum temperature of the stator coil is 110\u2103, which is located in the middle part of the end coil near the oil outlet, and the maximum temperature rise is 75\u00a0K. This is because of the inner thermal resistance and the long heat passage from the \u201ccool\u201d oil inlet to the \u201cwarm\u201d oil outlet. The calculated temperature of the air-core stator meets the class F insulation systems by the CFD analysis. The inner composite cylinder is eliminated to simplify the structural analysis and a parallel-face symmetry model of one slot air-core stator is built"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002077_j.mechmachtheory.2020.104209-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002077_j.mechmachtheory.2020.104209-Figure4-1.png",
        "caption": "Fig. 4. Force diagram of the i th limb assemblage with two passive joints.",
        "texts": [
            " According to the Newton\u2019s 2nd law, the equilibrium equation of the moving platform can be formulated as \u2212 3 \u2211 i =1 F Ai + F P = 0 ;\u2212 3 \u2211 i =1 r A i \u00d7 F Ai \u2212 3 \u2211 i =1 M Ai + M P = 0 (11) where r Ai denotes the position vector of passive joint A i measured in the frame O-xyz . The structural compliance of limb body is further included into the kinetostatic model by modeling it as an elastic spatial beam with a number of discrete units e j ( j = 1, 2,... N 0 ) through the finite element (FE) method [42] . Fig. 4 demonstrates the force diagram of an individual limb assemblage. As shown in Fig. 4 , the i th limb assemblage is meshed by Euler-Bernoulli beam element having bending, extending and torsional deflections. With the FE formulation, one may derive the elastic displacement of an individual limb measured in the frame A i -x i y i z i as the follow x i = k \u22121 i f i ( i = 1 , 2 , 3 ) (12) x i = [ ( x L Ai )T ( x A Ai )T ... ( x L Bi )T ( x A Bi )T ... ] T 6 N 0 +6 (13) f i = [ f T Ai m T Ai ... f T Bi m T Bi ... ]T 6 N 0 +6 (14) where x i , k i and f i are the elastic displacement coordinates vector, the stiffness matrix and the load vector of the i th limb assemblage, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002493_j.jnucmat.2020.152573-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002493_j.jnucmat.2020.152573-Figure2-1.png",
        "caption": "Fig. 2. (a) Morphologies of W-Cr coatings on the RAFM steel plates, (b) the scanning strategy in one layer of the LMD process.",
        "texts": [
            " The LMD system as a 1070 nm wavelength laser, a five-axis computer-controlled orkstation, and a chamber filled with high-purity argon where he oxygen content is controlled to be below 20 ppm in order o avoid oxidation. The mixed W-Cr powders were delivered to he molten pool through a coaxial nozzle at a feeding rate of 6- g/min. In order to release the thermal stress between the RAFM teel substrates and the deposited W-Cr coatings, the RAFM steel lates were heated under 400 \u00b0C for 30 minutes before deposition y a thermostatic heating platform. The heating was continued unil 30 minutes after the LMD process was finished. Five coating samples (see Fig. 2 a) with lengths of 20 mm, idths of 10mm, and heights of 2 mm were fabricated using differ- nt process parameters as listed in Table 2 . All the samples were repared using the same scanning strategy as shown in Fig. 2 b. or the multi-layer scanning, the starting point and the scanning aths are the same for each layer. In order to evaluate the effects f process parameters, a metric of surface energy density ( E s ) is ntroduced in this research by Eq. (1) . s = P dv (1) n which, P is the laser power, d is the hatch distance, and v is the canning speed. The calculated values of E s are listed in Table 2 . Microstructures on the cross-sections perpendicular to the laser canning direction (SD) were observed using a Nikon LV150 N ptical microscope (OM) and a Tescan Mira3 LMH field emision scanning electron microscope (SEM)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003484_9781119711230-Figure7.2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003484_9781119711230-Figure7.2-1.png",
        "caption": "Figure 7.2 Daksh remotely operated vehicles.",
        "texts": [
            " These defense robots have many functions, including: \u2022 Carrying heavy equipment \u2022 Rescuing wounded soldiers in combat zones \u2022 Operating in dangerous situations to keep soldiers at a safer distance Some field robots in defense applications are beginning to be outfitted with weapons for offensive capabilities. Defense robots are beginning to become a common part of military campaigns, assisting with maintaining the safety of troops and providing a tactical advantage in almost any combat situation. As the military keeps on testing and succeeding with various types of defense robots, the market is expected to see strong growth. The Indian Army has submitted a request for 200 Daksh remotely operated vehicles (ROVs) to defuse explosives (Figure 7.2). Furthermore, the Defence Acquisition Council (DAC) has endorsed approximately 544 robots for the Indian Army from indigenous source. The robots will be utilized for reconnaissance and can convey appropriate ammo [11]. 120 AI and IoT-Based Intelligent Automation in Robotics 1) Discriminating Between Targets In some situations it is extremely difficult to design a robot that can differentiate between a soldier and a civilian in order to apply the principles required by the Rules of Engagement (ROE) and Law of War (LOW), especially as agitators act like regular people"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000869_s11668-019-00593-2-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000869_s11668-019-00593-2-Figure2-1.png",
        "caption": "Fig. 2 Location of failure: (a) schematic showing arrangement of rolls in hot strip mill, (b) schematic of pinch rolls, (c) on-site image of location of failure",
        "texts": [
            " It goes through a secondary de-scaler to remove any scale that developed during the rolling process. This is followed by final thickness reduction in 4 high, 6 stand finishing mill. After attaining the desired reduction in thickness, steel sheets pass through a set of rolls and are subjected to laminar cooling. Here, water flow rate is controlled to attain the desired cooling rate and phase transformation. It then goes to the down coiler, and the finish rolling temperature (FRT) determines the final microstructure in the hot-rolled product. Figure 2a shows the schematic of the overall arrangement of rolls in hot strip mill. There are a number of rolls with specific functions. There are bend rolls, deflector rolls and entry rolls, and all of these fundamentally allow the workpiece to enter the forming roll in the desired alignment where reduction in thickness takes place. There are two sets of cradle rolls used for coiling of the steel sheets. Uncoiling takes place with the help of pinch rolls. There are top pinch rolls and bottom pinch rolls, and they work in conjunction to maintain the continuity of the hot rolling process as shown in Fig. 2b. These are driven by a motorconnected bearing assembly. Failure occurred in the bottom entry pinch roll in its drive side as shown in Fig. 2c. The bearing was found to be jammed and broken into multiple pieces. It failed after 5 months of service as against an expected life of at least 1 year. This reduction in service to less than half of its expected life is the motivation behind an in-depth failure analysis to understand its root cause and take corrective actions to prevent such failures in the future. Figure 3 shows a schematic of a pinch roll bearing used in HSM. It has three major elements, namely outer ring, inner ring and barrel-shaped rollers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000191_wcnc.2019.8885541-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000191_wcnc.2019.8885541-Figure6-1.png",
        "caption": "Fig. 6. COMSOL Simulation of Freelinc board coils.",
        "texts": [
            " From this figure, we can observe that within the transmission range the PDP is above 90% for a packet transmission rate of \u223c2000 packets/secs. As in our experiments, the packet size is set to 11 bytes and this provides a data rate of \u223c22 KBps which is a sufficient transmission rate for the typical sensing operation. As expected the PDP drops significantly with the increase in packet transmission rate. To compare the experimental results with the simulations, we imitate the Freelinc board in the COMSOL simulator. Fig. 6 shows the simulation environment. The lower coils are for the transmitter which consists of two ferrite-core and one aircore coils, whereas the receiver (the upper one) has air-core. We keep the dimensions identical to that of the Freelinc board as reported in Table I for accurate emulation. In particular, the twisted wire segments protruding from the ferrite core coils are exactly as used in Freelinc board and provide additional inductance and capacitance. COMSOL does simulate the impact these fine features due to its ability to use finite element models for simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002009_s42417-020-00259-6-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002009_s42417-020-00259-6-Figure8-1.png",
        "caption": "Fig. 8 Wind turbine test bench and fiber grating monitoring system",
        "texts": [
            "\u00a0(6), (7) and (8), the load of the j-th meshing gear tooth can be expressed as Since the load on the ring gear changes with time in actual work, the dynamic load can be expressed as Fq(t), then Eq.\u00a0(8) can be rewritten as where Fq(t), Fj(t), \u0394q(t) and \u0394j(t) are the total load and deformation of the ring gear and the load and deformation of the j-th meshing gear at time t. The dynamic load of the ring gear presents periodic fluctuation, and it is divided into n sub-steps. The deformation of each sub-step along the contact line of the meshing teeth is shown in Fig.\u00a08. According to Fig.\u00a07, when the jth tooth enters into meshing, the meshing position is located at the tooth root, and the meshing force is small. Its bending deformation and elastic deformation are small. With the revolution of the planet, the load on the jth tooth increases gradually, and the elastic deformation increases, and the bending deformation also presents an increasing trend from the root to the top of the tooth. In the same way, the elastic deformation and bending deformation of ith tooth decrease gradually when the ith tooth is out of meshing",
            " Compared with the radial direction, the circumferential direction of the ring gear side should be changed significantly, reaching the maximum value near the root circle and between the two adjacent teeth, and the measurement of the circumferential strain can better reflect the health state of the planetary gearbox. So, the FBG is arranged in the area with the maximum circumferential strain of the ring gear side, and the strain value of this part is more sensitive to the change of the state of the planetary gearbox. In this paper, the wind turbine test bed produced by SQI company is taken as the research object, as shown in Fig.\u00a08. The motor provides the input speed, and the magnetic powder brake provides the load of the planetary gearbox. The fiber grating used in the experiment is 2\u00a0mm in length and 0.125\u00a0mm in diameter. According to the above analysis, the fiber is pasted on the position of the tooth root at the middle side of the 6th and 7th tooth anticlockwise in the direction perpendicular to the ground, as shown in Fig.\u00a08c; the fiber is led out from the side of the planetary gearbox, as shown in Fig.\u00a08b; the dynamic monitoring system of the fiber grating is shown in Fig.\u00a08d. In the experimental data collection, FBG sensor is pasted on the position of the middle root of the 6th and 7th tooth of the counterclockwise (Y-axis) perpendicular to the ground, as shown in Fig.\u00a08b (circumferential direction). Therefore, this position should also be taken as the dynamic strain output point (as shown in Fig.\u00a03) when extracting the simulation results. And, the simulation results are, respectively, obtained under the condition that the motor output frequency is 7.5\u00a0Hz and the load torque is 59\u00a0N\u00b7m The strain time history of the ring gear root under the experimental conditions is shown in Fig.\u00a09. To effectively extract the time-domain strain history of the gear root in the working process, this paper uses NI USB 6009 acquisition card (sampling frequency: 5000\u00a0Hz, system detection accuracy: 4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002161_9783527813872-Figure11.10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002161_9783527813872-Figure11.10-1.png",
        "caption": "Figure 11.10 Rewritable memory based on the reversible conducting filament formation.",
        "texts": [
            " In nanoparticles or quantum dots containing composites composed of charge transporting phase separated by metal nanoparticles or by quantum dots, the transport is controlled by the tunneling through the nanoparticles (NPs) or quantum dots (QDs), which is limited when the localized energy levels are filled with charges. Such systems often show a negative differential resistance, i.e. decreasing current with increasing voltage in current\u2013voltage characteristics. Probably, the most advanced and ready-to-market solution of resistive memories is based on a reversible conducting filament formation in the insulating inorganic or polymer matrix (see Figure 11.10). The filaments may be formed by diffusion of metal atoms or clusters from tips-like structures formed during the top electrode deposition and the counter drift of indium from indium-tin-oxide (ITO) bottom layer as it was observed in the case of ReRAMs prepared from tris(8-hydroxyquinolie)aluminum (Alq3) [61], PMMA [62], plasma-polymerized styrene [63], or a commercially available n-type perylene derivative semiconductor N1400(Polyera) as resistive layers [64]. In these structures, the growth of the filaments was well documented by 3-D imaging using a combination of X-ray photoelectron spectroscopy (XPS) and time-of-flight secondary ion mass spectrometry (ToF-SIMS) depth profile analysis [63] and the switching mechanism was interpreted in terms of reversible activation/inactivation of the same conductive path"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003122_s11012-021-01410-7-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003122_s11012-021-01410-7-Figure4-1.png",
        "caption": "Fig. 4 Inter-shaft bearing\u2019s structure size and heat transfer network. a Structure size. b Heat transfer network",
        "texts": [
            " 3 The viscosity-temperature relationship of the lubricant Table 2 The experimental data of the kinematic viscosity at different temperatures [18] Temperature TL ( C) 30 40 50 60 70 80 Kinematic viscosity m (mm2/s) 15 10 7.8 5.9 5 4 2.3 Unsteady-state heat transfer model under the dynamic load The material of bearings is bearing steel GCr15, thus, Biot numbers of the outer ring, the inner ring and cylindrical rollers, are small enough (Bi\\0:1). The lumped parameter method, i.e. lumped heat capacity method [30], can be used for the heat transfer modelling for the bearing. The structure size and the heat transfer network of the inter-shaft bearing are shown in Fig. 4, where dL, d, di, Dm, do, D, dH are the diameters; B, dr, ar are the width, the diameter of the roller, the length of the roller; Tr, Ti, To, TLP, THP, TL, T? are the temperatures; Rri, Rro, Ri, Ro are the heat conduction thermal resistances; RLr, RLi, RLo, RLP, RHP are the heat convection thermal resistances. Assume that the temperature of the lubricant is TL \u00bc Tr\u00feTi\u00feTo 3 [11], and an average partition coefficient [11] is applied as follows: Qt \u00bc Qr \u00fe Qi \u00fe Qo; \u00f011a\u00de Qr \u00bc 0:5Qt; \u00f011b\u00de Qi \u00bc 0:25Qt; \u00f011c\u00de Qo \u00bc 0:25Qt; \u00f011d\u00de where Qr is the FHG distributed to rollers, Qi is the FHG distributed to the inner ring and Qo is the FHG distributed to the outer ring"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001776_tie.2020.3005108-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001776_tie.2020.3005108-Figure1-1.png",
        "caption": "Fig. 1. VFRM. (a) Model 1: 12s10r. (b) Model 2: 12s14r. (c) Model 3: 12s10r modular structure. (d) Model 4: 12s14r modular structure.",
        "texts": [
            " In Section IV, self- and mutual inductance of armature windings are calculated, for investigating the fault currents and magnetic isolation between phases, short circuit and open circuit fault are analyzed. In section V, the prototype is tested to verify the FE and theoretical analysis. Conclusion is given in Section VI. II. FEATURES OF NON-MODULAR AND MODULAR VRFM The 12-slot 10- and 14-rotor-pole VFRM have higher torque density and more sinusoidal back-EMF than 8-rotorpole machine, and have lower cogging torque, torque ripple and non-unbalanced magnetic force than 11- and 13-rotorpole VFRM [10]. The 12s10r and 12s14r VFRM are selected in this paper, as shown in Fig. 1 (a) and (b), and their modular structure are proposed as shown in Fig. 1 (c) and (d). All the machines have the same overall size, air gap length, stack length and shaft diameter, the difference between nonmodular and modular structure machines is that the nonmodular structures machines have field excited windings and armature windings on different stator tooth, while the modular structures machines share the same stator tooth. In addition, there is a flux gap exist between the stator modular structures. All the machines will be optimized by 2D finite element method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000130_coase.2019.8843076-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000130_coase.2019.8843076-Figure6-1.png",
        "caption": "Figure 6",
        "texts": [
            " In order to evaluate the control strategies, a test cube with an edge length of 40 mm has been welded. The tests were welded with the following parameters and boundary conditions: wire feed: 4 m/min; welding speed: 0.6 m/min; process modification: controlled short circuiting arc; filler material: 1.2 mm EN ISO 14341-A: G 3Si1; substrate plate: 3 mm S235JR; shielding gas: Ar: 98%, CO2: 2%; interpass temperature: 100 \u00b0C \u00b1 10 \u00b0C. The CAD model and the corresponding path planning for 27 layers without top and bottom layers, 50% infill and one wall layer are visible in Figure 6 and Figure 7. The following disturbance variables were deliberately provoked: The layer height and the bead width were chosen lower in the path planning than to be expected for the given parameter set. The actual volume build-up will thus exceed the parameterized layer height and the total expected volume. => ? Q\\^ Q> _`{\\ Q| } > Q` | > _`{\\ Q| } > } | > `\\^> the dimensions 10x10x40mm was connected to the cuboid. At this point, due to the incorrect parameterization of the bead height and width, larger material accumulations will occur especially on the disturbance cube"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003720_s0039-9140(96)02040-1-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003720_s0039-9140(96)02040-1-Figure7-1.png",
        "caption": "Fig. 7. Microcell with inverted microlithographicallly fabricated microelectrodes.",
        "texts": [
            " 6, (1) is the Pt contact, (2) is the mercury drop of 5/~1 volume, (3) is the sample of 10/tl volume, (4) is the capillary of the electrolytic bridge of the auxiliary and simultaneously of the reference electrode and (5) is the capillary for the inlet of nitrogen. In the microcell (Fig. 6), deaeration of the sample is carried out simultaneously with the pre-electrolysis (stripping voltammetry), but nitrogen is admitted higher than the sample level. In microcells [48, 49, 53], all three electrodes (working, auxiliary and reference) have been inverted, which allowed a decrease in sample volume, e.g. down to sub-microliter levels [53]. In the microcells shown in (Fig. 7) [48, 49], the three-electrode system was fabricated by microlithography. [48], Ag/AgC1 [49]). The sample volume is ~ 2/~1 [48,49]. The microcell shown in (Fig. 8) [53] consists of three strips of metal foil in a multi-decker sandwich heat-sealing film. In Fig. 8, (1) is the working microelectrode (4/zm Pt foil) and (2) is the auxiliary microelectrode (100 /lm Ag foil). This micro cell has a sample volume of 0.05-1.0/11. The use of a very small current in the microcells shown in Fig. 7 and 8 obviated the need to separate the electrode compartments by diaphragms. In thin-layer electrochemical cells [79,80] which will be discussed below, both types of working electrodes (usual and inverted states) have been applied. In such microcells the sample is usually in the form of a very thin layer between the working and auxiliary electrodes. Although thin-layer mi- crocells are applied for flow-through systems more often [54,55], there are also some thin-layer microcells [11,14,26,27,32,38,40,41] for a static sample that are of interest"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003659_mssp.1998.0190-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003659_mssp.1998.0190-Figure3-1.png",
        "caption": "Figure 3. The worst selection of measurement points; all measurement positions are on a straight line.",
        "texts": [
            "erence vectors is the most important factor in#uencing the condition number of [R p ]. Hence, in order to minimise the condition number of [R p ], the response measurement points should be selected so that the cross-product term of the di!erence vectors can be as large as possible. 4.2. THE WORST SELECTION OF MEASUREMENT POINTS Upon careful examination of equation (12), the worst selection of measurement points can be derived analytically. If all the selected measurement points are on a straight line as shown in Fig. 3, the directions of all the di!erence vectors are the same. In this case, the matrix [R p ] is identically singular, since the cross-product and scalar triple-product terms of equation (12) are zero. Therefore, the accelerations at the origin cannot be determined no matter how many measurement points are selected. 4.3. THE BEST SELECTION OF MEASUREMENT POINTS WHEN n\"3 By investigating equation (17), the best con\"guration of the three measurement points to minimise the condition number of [R p ] can be derived"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure26-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure26-1.png",
        "caption": "Fig. 26. Rigid Traction Coupling with Idling Drag Eliminator",
        "texts": [
            " When used on winches and rail vehicles, the rocking brake lends itself to abuse, as it is a convenient brake on the transmission, for which service it is not intended. Largely because of rapid resulting wear the design came into disfavour in a number of applications. Idling Drag Eliminator. This name describes another form of rocking brake where the effort required for backward rotation against the drag torque is derived from the idling power of the engine, and servo application is not necessary. This recent development is illustrated by Fig. 26, showing a traction coupling of 260 h.p. at 700 r.p.m., with an idling drag eliminator for use when directly coupled to a traction type gearbox, with baulking ring jaw clutches for the gear change, and open jaw clutches for the forward and reverse drive. It will be seen that the brake drum is connected to a sun pinion, meshing with pairs of planet wheels mounted on a planet carrier bolted to the casing, which rotates with the engine; a second sun wheel of slightly smaller diameter is keyed to the runner shaft",
            " In the example shown the brake is of the magnetic type used in a \u201cCotal\u201d gearbox, to which the coupling in question was applied, but a friction lining was introduced between the magnet surfaces to decrease the braking effect and to permit slip with little wear. I t was the intention to magnetize the brake and close the ring valve during each gear change so as to decrease the work done by the clutches and brakes in the epicyclic gearbox, but it was found experimentally that a sufficient momentary cushioning effect could be obtained from a plain traction coupling in the simpler way described below. The helix method of ring valve actuation can also be used in connexion with an epicyclic drag eliminator of the kind shown by Fig. 26, the object being to reduce the drag torque to a very low figure by closing the ring valve before the runner shaft is stopped and slow backward rotation takes place. The principal problem in the helix ring valve coupling is the cushioning of the inertia forces which come into play at the limits of travel of the ring-valve actuating sleeve. These forces are increased in the design with the idling drag eliminator added. Since the ring valve works in a chamber which is partially filled with oil, it is possible to secure a dashpot action at the limits of its travel, but this is an expedient, and the heiix ring valve coupling is not to be recommended if the simpler and cheaper traction coupling can reasonably be applied",
            " He understood that with this combination the plate clutch might be much smaller ; or in other words, the allowable pressure on the friction facing might be greater than with a normal friction clutch which had to take the full duty as the inherent flexibility of the fluid coupling relieved the momentary starting loads involved. Could the author indicate approximately (as a percentage) what extra pressure might be allowed on the facing of a single-plate clutch used in this manner as compared with one used without a fluid coupling ? 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from The device illustrated in Fig. 26, p. 118, embodying an epicyclic gear to give the same action as a rocking brake was ingenious, as also was its application to a creeping feed gear ; in the latter case it would appear that when feeding at one-tenth motor speed a considerable proportion of the total power of the motor was being dissipated, but there would be some increase in torque over the normal running torque. Could the author give an estimate of the percentage by which the normal running torque was thereby increased ? The application of the fluid coupling to a high-powered locomotive described in the latter part of the paper was particularly interesting and the use of the Legge synchro-coupling was another instance of the author\u2019s ingenuity in extending the application of the coupling and providing auxiliary mechanism which ensured success in its general performance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000035_s12555-018-0636-2-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000035_s12555-018-0636-2-Figure1-1.png",
        "caption": "Fig. 1. Definitions of the reference coordinate systems and vectors for spacecraft rendezvous and docking.",
        "texts": [
            " Body frame FT is considered to be coincide with the LVLH frame, i.e., xT is parallel to the vector rt and points to the radial direction; yT is along the opposite direction of docking port, and zT is perpendicular to the target orbit and three mutually perpendicular axes complete the right-hand system; similarly, in frame FP, yP points toward the docking port on the pursuer spacecraft and three mutually perpendicular axes coincident with the principle axis of inertia. All the frames are shown in Fig. 1, where rt is the inertial position of the target spacecraft represented in the frame FT , while r p is the inertial position of the target spacecraft represented in the frame FP. 2.1. Relative attitude dynamics Suppose the target and pursuer are rigid bodies, and the relative attitude of FP with respect to FT can be described by modified Rodrigues parameters (MRP) as [29] \u03c3 \u225c \u03b6 tan (\u03d1 4 ) , where \u03d1 \u2208 (\u22122\u03c0,2\u03c0) is the rotation angle about the Euler axis, and \u03b6 is the unit vector of the Euler axis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003645_s0166-1280(98)00321-2-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003645_s0166-1280(98)00321-2-Figure6-1.png",
        "caption": "Fig. 6. The decomposition where scission of the Ca\u2013O bond of the protonated isopropoxy group is rate-determining.",
        "texts": [
            " Models were assembled by one of the alkoxy groups generated from the primary (methanol, ethanol, n-propanol, n-butanol), the secondary (s-butanol, 2-propanol) or the tertiary alcohol (t-butanol), some Oss and one Si atom and were used for calculations of the activation energy for the decomposition of the alkoxy groups. For the decomposition of the alkoxy groups where the extraction of Hb by the oxygen atom of the group, as illustrated in Fig. 4 (hereafter the case of 2-propanol is shown as the example in Figs 4\u20136), is ratedetermining, the heat of formation was calculated in the same way. In order to examine the decomposition of the alkoxy group where scission of the Ca\u2013O bond (Fig. 5) or scission of the Ca\u2013O bond of the protonated alkoxy group (Fig. 6) is rate-determining, the heats of formation were calculated in the same way. In the latter case, the formal charge of the model was set to a unit positive charge because of the attached proton. Calculated results for the case where the dehydration occurs by scission of Ca\u2013O bond of the isopropoxy group on SiO2 are shown in Fig. 7. The heat of formation of the model was calculated when separating the Ca from the oxygen in the group at 0.1 A\u030a intervals from 1.2 to 4.0 A\u030a as optimizing the structure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000764_978-3-319-49058-8_68-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000764_978-3-319-49058-8_68-Figure4-1.png",
        "caption": "Fig. 4. Hydraulic system",
        "texts": [
            " All system elements move in a complete accordance with the system dynamics [17, 18]. In other words, in order to provide realistic behaviour of the virtual system, the real system is substituted with its mathematical model \u2013 the set of differential equations that describe its dynamics. In that way, integration of differential equations provides the information that real systems gets through appropriate sensors. The graphical representation of the equipment faithfully resembles the actual real world system. In the upper right corner of the Fig. 4 one could see the electrical cabinet (consisting of the control unit and low voltage electrical components). On the left side of Fig. 4, the working surface is placed with all its components (servo valve, hydraulic cylinder, linear position sensor, load, etc.). The entire virtual area is covered with 7 cameras, which allows for a user to see any part of the system that he is interested in. Cameras can move in all four directions (left, right, up and down), as well as to rotate or accelerate. Figure 5a shows the oil reservoir, electric motor and butterfly pump. Using the camera options (such as different angles, and zooming) can provide a closer look at the system components that might be of interest (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003768_1.2833790-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003768_1.2833790-Figure7-1.png",
        "caption": "Fig. 7, SPB has the largest maximum pressure followed by TPB, HB, and TFB. Although the maximum pressure of SPB is high, the peak of the pressure distribution of TPB in the circumferential direction is broader than that of SPB. As for the minimum negative pressure that is illustrated in Fig. 6, TFB has the smallest value and HB has the largest value, while SPB and TPB have the same value. The depth of the separating groove was assumed to be 80 /j^tn. If the depth of the separating groove is deepened, the negative pressure will be smaller. For HB, the negative pressure is generated near",
        "texts": [
            "2 Pressure Distribution and Dam Effect. The pres sure distribution of a hydrodynamic fluid bearing is one of the most important performance indexes. In order to improve the load capacity and the stiffness of the bearings, it is im portant that only high positive pressure is generated in the bearings. In particular, it is critical to decrease disturbances by eliminating the negative pressure. Figure 6 shows the overall pressure distribution of the four types of bearings under maxi mum axial stiffness k.^. condition. Figure 7(a) shows the pres sure distribution along the circumference of the turning radius for HB and along the circumferences with mean radii for SPB, TPB, and TFB. Figure 1(b) shows the pressure distribution along the boundary lines between the steps and the lands for SPB and HB and the pressure distribution along the boundary lines between the tapers and the lands for TPB and TFB in the radial direction. The ordinate in Fig. 6 and Fig. 7 is defined as P = (p - p\u201e)/pa. From the maximum pressure illustrated in Fig. 6 and in 118 / Vol. 121, JANUARY 1999 Transactions of the ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 02/05/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use the inner and outer boundaries. This is inevitable because the pumping action in these regions is small and it can be a main source of disturbances. Figure l{h) indicates that the pressure distribution within the pocket of SPB is kept almost constant due to the dam effect. However, the pressure peak near the center of TFB is relatively low because there are no dams at either edges that can reduce side-flow. Therefore, side-flow can be reduced and pressure can be increased by adding dams at both the inner and outer sides of the taper. This can be illustrated by comparing the pressure distribution of TPB and TFB in Fig. 6(^7), {d), and Fig. 7(fc). In Fig. 7(i>), the pressure on the boundary lines between the taper and the land of TPB is smaller than that of SPB is because the effects of the dams near the boundary lines become smaller and the bearing characteristics in these regions becomes similar to that of TFB. 4.3 Comparison of Load Capacity of Double-Sided Thrust Bearings. In general, a disk spindle rotor is supported by a double-sided thrust bearing in the axial direction, upper and lower, as is shown in Fig. 8. Force F, that is generated in the lubricant film at the lower side acts on the rotor downwards and force F\u201e that is generated at the upper side acts upwards"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003066_j.compeleceng.2021.107267-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003066_j.compeleceng.2021.107267-Figure15-1.png",
        "caption": "Fig. 15. Exploded view of the experimental testbed assembly.",
        "texts": [
            " 14, where the implementation of simulation tools on the design proposed by the student can be observed in greater detail, where, prior to the implementation stage, the performance simulation was carried out of the prototype against possible falls and blows in the critical points of the drone. Moreover, the base of the V-Model from Fig. 13 shows the experimental quadcopter prototype the student deployed; which was later was tested and the controller outputs of the implemented fuzzy controller is shown in stage 5, in addition to the validation of the system that was performed by the student by plotting the error through time of the stabilization signal. Therefore, Fig. 15 highlights the most relevant elements of the implemented testbed from the student, shown through an exploded view in order to clarify the interaction between those elements. Thus, the green bidirectional arrow refers to the exchange of information between the control unit where the fuzzy logic controller is implemented (black PCB) and the board containing the information of the position sensor, velocity and acceleration of the prototype\u2019s center of mass (red PCB). In the same way, the red arrows refer to the interaction of each driver of the Brushless DC (BLDC) motors with the card that controls the speed of the motors (red PCB)",
            " 6 and 7), starting with greater importance at the beginning of the project where the concepts are presented, going through almost no participation in the advisory role where students take the leading role, until the ending point where the teacher takes the primary role again to validate and give feedback to the students on the development of the project. From Fig. 14, it is clear that the student could develop the designing skills of a mechanical structure in spite of the electronics engineering background of the testing subject, which clearly validates that the student learned new competences through the experimental development of the quadcopter prototype. Yet, there were some limitations in the mechanical design shown in Fig. 15, since the student\u2019s inexperience in the field, made him had some limitations on some finer design details, such as: the selection of the structural materials, the total weight of the prototype and finally its aerodynamics. As a consequence, the final prototype had low energetic efficiency due to the high power required to achieve the steady state in the air, due to the high weight of the quadcopter\u2019s structure. Therefore, as stated before, the authors are planing as future work to validate the proposal through the study of a pilot multidisciplinary group, seeking to enable a statistical analysis of the competences developed through the V-model teaching process, which would also presumably allow to enhance different disciplinary competences through the mutual collaboration between the students in the developing process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure31-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure31-1.png",
        "caption": "Fig. 31. Rigid Type of Traction Coupling",
        "texts": [
            " A variety of solutions were considered, one successful method being the use of an intermediate mounting ring of relatively slender section bolted to the impeller at six equidistant points and to the driving disk at six points spaced midway 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from between them. In another case tangential links were used, each being bolted at the ends to the impeller and at the,mid point to the driving disk: although their strength was far in excess of the driving torque, they could not withstand the stresses arising when the engine was running through a torsional critical speed. The best practice at present is illustrated by Fig. 31, from which it will be seen that the coupling is 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 127 registered for concentric running by a spigot of small diameter in the bore of the driving disk, and the bolts fitted in the impeller driving flange pass through clearance holes in the disk to permit of creep due to differential expansion, heavy Grover washers being used to give additional resilience to the clamped joint. the overhanging weight of the rigid type of traction coupling, Fig",
            " 34 is a chart showing nine convenient schemes that may be used, the points of which are explained by the notes on each diagram. The driving of railcar auxiliaries, such as the generator and compressor, often presents a problem when space is limited, particularly since the output shaft of the coupling is the most convenient point to take a V-belt drive; yet it is undesirable to impose the belt pull on the overhung coupling, and an outboard bearing is attended-by the difficulties already described when used with a rigidly mounted coupling as in Fig. 31. 9 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 130 PROBLEMS OF FLUID COUPLINGS FLUID COUTLING FLEXIBLE COUPLING \"SILENTBLOC\" TYPE MAY BE USED TO TRANSFER PART WEIGHT a. Rigid traction weight of complete engine bearings. FLUID Ey OUTBOARD BEARING SELF-ALIGNING TYPE INNER RACE MOUNTED O N \"SILENTBLOC'BUSH coupling. coupling ca Assume irried by b. Rigid traction coupling with resiliently mounted outboard bearing. For use when weight of coupling is too great to be carried by engine bearings",
            " Heavy sugar moulds were carried on a series of turntables attached to a chain, moving along slowly at a speed varying from 3 to 53 ft. per min. On account of the very slow speed the power needed to drive the loaded conveyer was small, although the moving mass amounted to 240 tons. The total ratio of reduction of the combined gears was no less than 3,00011. A traction-type fluid coupling was fitted to the spindle of a direct-current motor of 15 h.p. running at 720 r.p.m., the type of coupling being shown in Fig. 31, p. 126. The coupling would normally transmit a torque of 73 1b.-ft., and it was designed to stall against a torque of 137 1b.-ft. at full speed, thus effectively protecting a conveyer costing over Ll8,OOO against the possibility of a destructive overload. Mr. C . E. FAIRBURN (London, Midland and Scottish Railway) said that in the last few years it had been his good fortune to see many Diesel railcars in the United States and in Europe, and he thought it was now fairly clear that with an engine of below 300 h"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001079_1464419319862456-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001079_1464419319862456-Figure4-1.png",
        "caption": "Figure 4. Schematic diagram of test bench.",
        "texts": [
            " When the signs change again, it indicates that the solution of the system enters the stable stage once again. In this chapter, the rotor system with bolted flange joints is taken as the experimental object. The experimental results are compared with the simulation results to verify the correctness of the modeling method of rotor system with bolted flange joints and the analysis of its motion evolution law, and also verify the feasibility of tracking the amplitude\u2013frequency curve of rotor system by IHB\u2013AFT method. The experimental arrangement is shown in Figure 4. The two ends of the spindle of the rotor system are supported on rolling bearings in two bearing pedestals. The parameters of the rolling bearings are shown in Table 1. The bearing pedestals are fixed on the foundation platform. The rotor system is driven by the motor at the right end. The basic parameters of the rotor system and the parameters of bolted flange joints are shown in Table 2. The test equipment used in the experiment includes CA-YD-502 eddy current sensor, power amplifier, and NI9229 signal acquisition module"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000655_s00170-016-9183-2-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000655_s00170-016-9183-2-Figure11-1.png",
        "caption": "Fig. 11 Percentage impact of inputs on a deformation energy and b unclamped geometric accuracy",
        "texts": [
            " Figure 10a shows that the combined effect of step down and sheet thickness produces the highest variations (number of peaks) in the deformation energy of the pyramid ISFformed part followed by the combined effect of the input pairs (step down and wall angle and wall angle and sheet thickness). Figure 10b shows that the combined effect of pair sheet thickness and wall angle produces the highest variations in the unclamped geometric accuracy of the pyramid ISF-formed part followed by the combined effect of pairs (tool diameter and wall angle and tool diameter and sheet thickness). Sensitivity of the inputs to the two characteristics was computed based on finding the maximum/minimum and observing the number of peaks from the 2D and 3D plots, respectively. It is found from Fig. 11a, b that the sheet thickness influenced deformation energy the most followed by step down, wall angle, and tool diameter. For unclamped geometric accuracy, wall angle has the highest influence followed by sheet thickness, tool diameter, and step down. Overall, sheet thickness plays a greater role and has the highest influence for both response characteristics. The analysis suggests that the sheet thickness value should be low to achieve lower deformation energy and higher to achieve higher unclamped geometric accuracy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002020_s12206-020-1030-6-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002020_s12206-020-1030-6-Figure12-1.png",
        "caption": "Fig. 12. Experimental platform for object grasping.",
        "texts": [
            " Based on the motion characteristic of the finger mechanism, a prototype model is developed to investigate the effectiveness of the finger mechanism in bidirectional object grasping, as depicted in Fig. 11. The presented finger mechanism has a parameter set of 1 2 5l l l= = = 5.5 cm, 3 4l l= = 2 cm. The link AB and CD is designed as S type to avoid motion interference and obtain a maximum grasping range. In Fig. 11, the yellow transparent area denotes the overall grasping range of the finger mechanism. To achieve object grasping, two modular fingers have been fabricated and form a configurable gripper, as depicted in Fig. 12. Each finger mechanism is actuated through a closed loop tendon. And the two tendons are actuated by the wire wheel which mounted on the output shaft of the micro servo motor. In the present configuration, the gripper could grasp object from outside to inside as the wire wheel rotates clockwise. Similarly, the gripper will grasp object from inside to outside as the wire wheel rotates counterclockwise. Two groups of grasping experiments have been conducted to investigate the effectiveness of the finger mechanism in bidirectional object grasping"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure8-1.png",
        "caption": "Fig. 8. Instrumented torsion platform.",
        "texts": [
            " The method can be used with the measurements collected using the torsion platform described in Section 4, or another equivalent platform. The method requires a set of estimates of the mass moment of inertia of the object about a vertical axis (IRzz, defined in this section). In the case of a torsion platform, these are obtained by measuring the period of oscillation of the base platform with the object mounted in a series of specified configurations. The proposed method considers an object (Fig. 4) with its CM located at rC with respect to frame O. Frame O has the axis zO aligned with the axis of rotation of the system (Fig. 8). Frame B is a body fixed frame located at an arbitrary location rB with respect to frame O and is initially parallel to it. Frame C is a body fixed frame with origin at the CM and parallel to frame B. The object has an unknown inertia tensor I with respect to frame C: I \u00bc Ixx Ixy Ixz Ixy Iyy Iyz Ixz Iyz Izz 0 B@ 1 CA \u00f010\u00de Vector rB can be measured directly: rB \u00bc xB; yB; zB\u00f0 \u00deT \u00f011\u00de while the vectors rC=B and rC are unknown: rC=B \u00bc xC=B; yC=B; zC=B T \u00f012\u00de rC \u00bc rB \u00fe rC=B \u00bc xC ; yC ; zC\u00f0 \u00deT \u00f013\u00de The inertia of the object prior to any rotations can be determined relative to frame O from the parallel axis theorem: IO \u00bc I \u00fem rTCrCI3 rCrTC \u00f014\u00de where I3 is a 3 by 3 identity matrix, and m is the mass of the object",
            " The errors between the results from SolidWorks and our equations differ by a maximum error with order of magnitude of 1 10 8, which is expected because the inertias from the computer-generated model are reported to eight decimal points. The inertia tensor and center of mass were calculated to be: The differences between the SolidWorks-reported inertia tensor and CM position vector and our calculated values are: g b a xB yB zB An instrumented torsion platform was developed and built to experimentally validate the proposed method in Section 3 (Fig. 8). In this section, we describe the design, equations of motion, calibration and uncertainty analysis of this apparatus. An instrumented torsion platform prototype was built and is shown in Fig. 9. The system is designed to oscillate a small object about a vertical axis. The rotation shaft is supported by a small double bearing block that is attached to an aluminum fixed frame, as shown in Fig. 8. The top end of the shaft is rigidly attached to a rotation plate with a clamping hub. The shaft is stepped and rests on the edge of the bearing to carry the weight of the plate. The bottom end of the shaft is connected to the top end of a linear torsion spring. The bottom end of the spring is connected to the fixed frame with a rigid coupler. The spring is made from 316 stainless steel round wire with a diameter of 2.29 mm. The diameter and length of the spring are 38 mm and 60 mm, respectively",
            " For a system with negligible friction and an oscillation period on the order of 1 s, the uncertainty of uI is calculated to be on the order of 1 10 5 kg m2 (Section 4.7). The experimental procedure to determine IOzz using our apparatus is outlined below. 1. Calibrate the torsion spring for positive angular displacements (Section 4.4). 2. Configure the electronics to record the half-period of oscillation during the positive displacement part of the motion. 3. Rigidly attach the rotation plate, attachment blocks, and test specimens in the desired position and orientation, to the torsion platform (Fig. 8). 4. Collect dataset: (a) Start the motion of the platform by providing an initial angular displacement h0 and releasing it from rest. A h0 of approximately p=3 is recommended. (b) Record the amplitudes and half-periods of successive oscillations for the positive displacement part of the motion until the amplitudes have reduced by approximately 50%. (c) Calculate Tk, the period of oscillation, assuming that it is twice the half-period recorded. (d) Report any external factors that could have affected the dataset obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003483_s40964-021-00197-z-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003483_s40964-021-00197-z-Figure5-1.png",
        "caption": "Fig. 5 Detail 1 of Fig.\u00a0 4: geometric parametrization of the (SBP) scanning strategy",
        "texts": [],
        "surrounding_texts": [
            "bles Figure\u00a06 shows the required characteristic points, vectors ( \ufffd\u20d7v and \ufffd\u20d7n ) and bisector (B) that are used in the analytical modeling of the scan distance series CH . CH is considered as a characteristic distance of the problem since it describes the series of active scanning displacements in contrast to the jump displacements that are performed along the hatch space jumps. Once the distance CH expression is formulated, the calculation of the total productive scanning length LP of the areas A and B of the rectangle is straightforward; (1) 1L1 = L2\u22152, (1.1) 2L1 = L1 \u2212 L2, (1.2) . 1 3 the total productive length corresponds to the sum of the series of CHi as expressed by the Eq. (2): Figures\u00a06 and 7 show the geometrical parameters that allow the calculation of the distances CHi as follows: \u2022 The parameters and notations utilized are: \u2022 ( xC, yC ) and ( xH , yH ) \u2236 resp. are the coordinates of the points C and H; (2)LPZone A or B = \u2211 i\u2208{index set} CHi, \u2022 \ufffd\u20d7n : normal unit vector to the skeleton at any point H; it corresponds the unitary vector of the line (CH); \u2022 \ufffd\u20d7v : carrier vector of the bisector (B) that is perpendicular to the vector \ufffd\u20d7n; \u2022 The bisector at the upper-left vertex of the rectangle is denoted (B). Its equation is written as: According to the axis convention of Fig.\u00a04, it is easy to argue that: The modeling procedure of the distances CHi begins with the decomposition of the first quadrant related to the left upper vertex as shown in the Fig.\u00a07. Two horizontal axes are shown in this figure: \u2013 the first (upper) horizontal axis describes the true abscissa x (control abscissa); \u2013 the second (lower) horizontal axis describes the index abscissa k. Figure\u00a07 shows the characteristic abscissa that will help in the formulation of the total scanning length of region A."
        ]
    },
    {
        "image_filename": "designv11_14_0002434_j.procir.2020.02.184-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002434_j.procir.2020.02.184-Figure1-1.png",
        "caption": "Fig. 1: (a) Designed CAD model of the Barrel finishing machine; (b) developed setup of Barrel finishing machine: 1. barrel, 2. barrel protective strap, 3. universal coupler, 4. power shaft, 6. bearings, 7. motor, 8. inclined bed, 9. base structure, 10. wheels supporting columns, 11 supporting wheels, 12. variable frequency drive; (c) front view of the barrel showing additional strips.",
        "texts": [
            " It should support the process to establish it as an easy, economical and shape independent technique for simultaneous removal of oxides and minimisation of surface roughness. \u2022 Finish the corroded samples to find out the optimum value of processing parameters such as finishing time, rotation speed, the concentration of abrasive particles, and wet or dry finishing. \u2022 To find the effect of abrasive mesh (size) on finishing time. The barrel finishing setup has been designed and developed in-house. An attempt has been made to make the design simple, easy to manufacture, mechanically balanced and cost-effective. Fig. 1a and 1b show the CAD model and pictorial view of the setup, respectively. The mentioned nomenclature represents the various components utilised for the fabrication. It consists of a rotary barrel which was inclined at 45\u00b0. The barrel was inclined to take advantage of both the actions of forces, i.e. generated due to avalanche motion and centrifugal forces in case of horizontal and vertical barrel position, respectively. Three metallic strips were welded on the inner periphery of the barrel at 45\u00b0 inclination and 120\u00b0 apart from each other (Fig. 1c). It was adapted to promote the avalanche motion of the abrasives during motion and to prevent the accumulation of the abrasives on the internal wall due to centrifugal force. A threephase AC motor (Make: ABB, Model: M2BAX80MB4 IE2, 0.75 kW, 1425 rpm) was used to provide the rotary motion to the barrel. It was connected to the barrel via a power shaft supported by two bearings. The rotation speed of the motor was controlled by a variable frequency drive (VFD) (Emotron VSB AC drive, 0-600 Hz). All the design considerations and constraints are presented in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000764_978-3-319-49058-8_68-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000764_978-3-319-49058-8_68-Figure6-1.png",
        "caption": "Fig. 6. Hydraulic cylinder (left), uncovered cylinder (right)",
        "texts": [
            " The entire virtual area is covered with 7 cameras, which allows for a user to see any part of the system that he is interested in. Cameras can move in all four directions (left, right, up and down), as well as to rotate or accelerate. Figure 5a shows the oil reservoir, electric motor and butterfly pump. Using the camera options (such as different angles, and zooming) can provide a closer look at the system components that might be of interest (Fig. 5b). The user can remove the covers and have an inside view in internal structures of some system elements (Fig. 6), thus having a possibility to obtain deeper knowledge about the role and working principles of the particular part of the mechatronic system. Figure 6b shows different pressure levels that are highlighted in different colours. STE sciences are yet to implement all the potentials of distance learning. The main constraint is represented by the fact that these sciences require laboratory exercises. Software based virtual laboratories have emerged as a solution, and shown great potential in teaching and learning. Although the final goal of the virtual lab is to replace the real one, it can have another important role \u2013 serving for training before the experimentation in the real laboratory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000222_0954406219885979-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000222_0954406219885979-Figure3-1.png",
        "caption": "Figure 3. Relative displacements of the ball center and curvature centers at angular position qj due to dynamic effects.",
        "texts": [
            " Their final positions due to external forces can be calculated by substituting Ri cos qj Ri sin qj 0 T into equation (2), which is as follows xiqj yiqj ziqj 2 64 3 75 \u00bc Ri cos qj cos y \u00fe 1\u00f0 \u00de qLi 2 sin y \u00fe x 1\u00f0 \u00deqRi cos qj sin x sin y \u00fe Ri sin qj cos x \u00fe 1\u00f0 \u00de q 1\u00f0 \u00deq\u00fe1 Li 2 sin x cos y \u00fe y 1\u00f0 \u00deq\u00fe1Ri cos qj cos x sin y \u00fe Ri sin qj sin x \u00fe 1\u00f0 \u00de q 1\u00f0 \u00deq Li 2 cos x cos y 1 \u00fe z 2 6666666666666666664 3 7777777777777777775 \u00f03\u00de The coordinate of the outer race groove curvature center corresponding to the jth ball of the qth row is: 0 0 Li 2 T . Figure 3 illustrates the relative dis- placements of the ball center and curvature centers at angular position qj due to external loads. According to the Pythagorean theorem, the displacement equations for the ball are as follows R2 bqj \u00fe Z2 bqj ro ur 0:5D\u00fe oqj 2 \u00bc 0 \u00f04\u00de Riqj Rbqj 2 \u00fe Ziqj Zbqj 2 ri 0:5D\u00fe iqj 2 \u00bc 0 \u00f05\u00de where iqj and oqj denote the inner and outer contact deformations respectively;Rbqj and Zbqj denote the radial and axial distance between the ball center and the outer raceway groove curvature center respectively; Riqj and Ziqj denote the radial and axial distance between the inner and outer raceway groove curvature centers respectively, which can be calculated by using the following equations Riqj \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2iqj \u00fe y2iqj q \u00f06\u00de Ziqj \u00bc ziqj \u00fe Li 2 \u00f07\u00de Load equilibrium for the ball and the inner ring The force analysis of the ball in quasi-static equilibrium state is shown in Figure 4",
            " The constraint equations at xq yq zq 1 2 6664 3 7775 \u00bc cos y 0 1\u00f0 \u00deq sin y 1\u00f0 \u00deq Li 2 sin y \u00fe x 1\u00f0 \u00deq sin x sin y cos x sin x cos y 1\u00f0 \u00deq 1\u00f0 \u00deq\u00fe1 Li 2 sin x cos y \u00fe y 1\u00f0 \u00deq\u00fe1 cos x sin y sin x cos x cos y 1\u00f0 \u00deq 1\u00f0 \u00deq Li 2 cos x cos y 1 \u00fe z 0 0 0 1 2 66664 3 77775 x0q y0q z0q 1 2 6664 3 7775 \u00f02\u00de the ball are2,26 Qiqj sin iqj Qoqj sin oqj \u00fe 2Mgqj D cos oqj \u00bc 0 \u00f08\u00de Qiqj cos iqj Qoqj cos oqj 2Mgqj D sin oqj \u00fe Fcqj \u00bc 0 \u00f09\u00de where Qiqj and Qoqj denote the inner and outer contact forces respectively; Fcqj and Mgqj denote the centrifugal force and the gyroscopic moment at the ball; iqj and oqj denote the inner and outer contact angles respectively, which can be obtained from Figure 3 iqj \u00bc cos 1 Riqj Rbqj ri 0:5D\u00fe iqj \u00f010\u00de oqj \u00bc cos 1 Rbqj ro ur 0:5D\u00fe oqj \u00f011\u00de The contact forces can be calculated by using the Hertzian contact theory,2 and the centrifugal force and the gyroscopic moment can be expressed as follows Fcqj \u00bc 1 2 mdm! 2 !mqj ! 2 \u00f012\u00de Mgqj \u00bc J !Rqj ! !mqj ! !2 sin qj \u00f013\u00de where m and J denote the mass and Moment of inertia of the ball, respectively; dm denote the pitch diameter of the double-row self-aligning ball bearing; !, !mqj, and !Rqj denote angular velocity of the inner ring, orbital angular velocity of the ball and angular velocity of the ball about its own axis respectively; qj denotes the pitch angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure9-1.png",
        "caption": "Fig. 9. Torsion pendulum apparatus.",
        "texts": [
            " The inertia tensor and center of mass were calculated to be: The differences between the SolidWorks-reported inertia tensor and CM position vector and our calculated values are: g b a xB yB zB An instrumented torsion platform was developed and built to experimentally validate the proposed method in Section 3 (Fig. 8). In this section, we describe the design, equations of motion, calibration and uncertainty analysis of this apparatus. An instrumented torsion platform prototype was built and is shown in Fig. 9. The system is designed to oscillate a small object about a vertical axis. The rotation shaft is supported by a small double bearing block that is attached to an aluminum fixed frame, as shown in Fig. 8. The top end of the shaft is rigidly attached to a rotation plate with a clamping hub. The shaft is stepped and rests on the edge of the bearing to carry the weight of the plate. The bottom end of the shaft is connected to the top end of a linear torsion spring. The bottom end of the spring is connected to the fixed frame with a rigid coupler",
            "87041 kg and it was measured to be 0.870 kg on a calibrated balance. The components can be expected to have constant densities, therefore, the inertia tensors are expected to be in close agreement. Six different configurations were selected and ten datasets were collected for each configuration following the procedure in Section 4.6. The system was configured by connecting the shaft to the rotating plate in six different locations. The plate was displaced laterally in the directions of xO and yO in Fig. 9, and attached using the hole patterns near its center. The displacements used for each configuration are presented in Table 8. The SolidWorks model was set to match the experimental configurations of Table 8. The attachment blocks were removed for these validation tests. The oscillation period and decaying amplitude measurements for 10 datasets in configuration 1 are presented in Figs. 14 and 15 respectively. The bounded region on the plots represents the subset of measurements selected for further analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001499_s00366-020-00963-7-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001499_s00366-020-00963-7-Figure4-1.png",
        "caption": "Fig. 4 3D model of 1-DOF compliant mechanism",
        "texts": [
            " Flexure hinges B are connected with a lever amplifier with dimensions length L3 and L4, and thickness T3 and T4, respectively. The output motions of the lever amplifiers are focused to transform to the output motion of the OCM. Two sides of output of the OCM are mapped with two flexure hinges C with length L5 and thickness T5. Overall the mechanism is fixed at five supports by using screws. Fig. 3 Schematics of 1-DOF compliant mechanism 1 3 Out-of-fabrication-plane thickness of the OCM is labeled as W, as depicted in Fig.\u00a04. As described in Figs.\u00a03 and 4, the geometrical factors and their dimensions of the proposed OCM are given in Table\u00a01. According to the designer\u2019s experiences and expert\u2019s knowledge in the compliant mechanism, the most important design variables include length and thickness of flexure hinges. Behaviors of compliant mechanisms are strongly dependent on a change in cross-section of flexure hinges according to mechanic theory of deformation. Meanwhile, other parameters, e.g., angle between rhombic link and ground, \u03b1, is also sensitive to statics and dynamics of the proposed mechanism but influence or contribution of this angle on the performances is small"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure6-1.png",
        "caption": "Fig. 6. Vector-loop closures in 3-R(RRR)R+R HAM.",
        "texts": [
            " we can obtain the following Mm_i = AdT g Mm(i = 1, 2, 3) (16) The actuation joints of the antenna mechanism include three revolute joints Pi (i = 1, 2, 3) connected with the fixed platform and the revolute joint Q4 of the polarization device. For the inverse kinematics analysis, assuming that the position matrix \u25e6r and the attitude matrix \u25e6RC of the 3-R(RRR)R parallel mechanism platform are known conditions, the inverse kinematics problem is to obtain the actuation joint variables. The vector-loop closures in the mechanism are shown in Fig. 6. The loop-closure equation for each actuated limb is written as follows: \u25e6Pi + \u25e6Pi Gi + \u25e6Gi Qi = \u25e6r + \u25e6Ci Qi = \u25e6r + \u25e6RC C Ci Qi (17) where i = 1, 2, 3, \u25e6Pi Gi is the vector from Pi to Gi in the fixed coordinate system, \u25e6Gi Qi is the vector from Gi to Qi in the fixed coordinate system, \u25e6Ci Qi is the vector from Ci to Qi in the fixed coordinate system, C Ci Qi is the vector from Ci to Qi in the moving coordinate system. The loop-closure Eq. (18) can be written as follows: \u25e6Gi Qi = \u25e6r + \u25e6RC C Ci Qi \u2212 \u25e6Pi \u2212 \u25e6Pi Gi (18) The position vector of the moving coordinate system is \u25e6r = [r sin \u03ba cos \u03c2, r sin \u03ba sin \u03c2, r cos \u03ba]T , and the attitude matrix \u25e6RC of the moving coordinate system is as follows: \u25e6 RC = \u239b \u239c\u239d s2\u03c2\u03bd\u03b4 + c\u03b4 \u2212s\u03c2c\u03c2\u03bd\u03b4 c\u03c2s\u03b4 \u2212s\u03c2c\u03c2\u03bd\u03b4 c2\u03c2\u03bd\u03b4 + c\u03b4 s\u03c2s\u03b4 \u2212c\u03c2s\u03b4 \u2212s\u03c2s\u03b4 c\u03b4 \u239e \u239f\u23a0 (19) where s\u03c2 = sin \u03c2 , c\u03c2 = cos \u03c2 , \u03bd\u03b4 = 1 \u2212 cos \u03b4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000683_tmag.2016.2601886-Figure23-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000683_tmag.2016.2601886-Figure23-1.png",
        "caption": "Fig. 23. Prototype of 12S/10R PS-PMM-I.",
        "texts": [
            " Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. still exhibit larger average torque than PS-PMM-Is over the whole copper loss range. Further, among the four stator/rotor pole number combinations, PS-PMM-Is and PS-PMM-IIs with 11-pole rotor still have the optimal torque capability over the whole copper loss range. Prototype machine of 12S/10R PS-PMM-I is made to validate the previous analyses, as shown in Fig. 23. To reduce the cost, this prototype machine is assembled by existing inner TABLE VI PARAMETERS OF PROTOTYPE MACHINE (12S/10R PS-PMM-I) Parameter PS-PMM-I Parameter PS-PMM-I LAG (mm) 0.5 TOSTTO( mm) 1 LAA (mm) 25 TOSY (mm) 3 ROSO (mm) 45 \u03b8ROP (\u00b0) 18 ROSI (mm) 31.75 \u03b8RIP (\u00b0) 25.2 RISO (mm) 25.75 TRR (mm) 5 RISI (mm) 10.4 TPM (mm) 4 \u03b8OSTB (\u00b0) 8.12 \u03b8PM (\u00b0) 30 \u03b8OSTT (\u00b0) 4.94 TBRI (mm) 0.5 TOSTTB (mm) 3.0 stator (12 PMs), outer stator (12-pole) and modular rotor (10- pole). Hence, the parameters of prototype machine as shown in Table VI are different from the previous globally optimized parameters as shown in Table I. Moreover, for easing the fabrication, 10-pole modular rotor is mechanically connected by the lamination bridges (TBRI = 0.5mm) in the inner side, as show in Fig. 23 (d) and Fig. 24. Fig. 25 shows the predicted and measured phase backEMFs at rated speed (400rpm). The measured fundamental value is ~4% less than the prediction, which is mainly due to the end effect in 25mm stack length machine. The predicted and measured open-circuit cogging torque waveforms are shown in Fig. 26. It can be seen that the measured peak to peak value is slightly larger than the FE prediction. This difference is acceptable when considering the measurement error and assembling tolerance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000464_j.isatra.2015.12.017-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000464_j.isatra.2015.12.017-Figure9-1.png",
        "caption": "Fig. 9. The rotor AMB test rig.",
        "texts": [
            " In order to make sure the assumption is legit, we performed the same simulation for 2100 r/min (35 Hz) and the results are shown in Fig. 8, where the same simulation conclusion can be drawn. Specifically, for a good rolling element bearing, electromagnetic force does not increase the vibration; however, for an outer race defect, it can amplify the vibration clearly. 4. Experimental verification The experimental test rig for this study is a rotor AMB system designed and built as a research platform at Nanjing University of Aeronautics and Astronautics, as shown in Fig. 9. The rotor is supported by two radial and two thrust AMBs and is designed to operate at a maximum speed of 60,000 r/min. For this study, the rotor is supported by rolling element bearings rather than AMBs and AMBs are employed as the non-contact exciters to apply online electromagnetic force during the operation. All the electronic equipments employed to perform the test are shown in Fig. 10. One acceleration sensor is installed on the pedestal vertically as shown in Fig. 9. The control box is employed in this experiment to apply constant current to the AMBs, which generate vertical electromagnetic force on the rotor. Electric discharge machine is employed to create the incipient defect on the bearing outer race and the defect creating process is illustrated in Fig. 11(a). Fig. 11(b) shows the roller element bearing with 0.2 mm width defect on its outer race. The experiment is performed at 1500 r/min (25 Hz) under radial AMB force with 1 A bias current. The bearings are positioned such that the outer race defects are always in the loaded zone where a fault would most likely occur [26]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure11-1.png",
        "caption": "Fig. 11. Russian Scoop- Tube Coupling",
        "texts": [
            " Circulation can be maintained through a fluid coupling by means of an external pump, in which case the liquid discharged from the working circuit is collected in a casing and drains to a sump, its energy of discharge being lost. If a scoop tube instead of a pump be used to maintain the circulation, the small power required to drive the pump is saved, since the scoop tube utilizes the kinetic energy of the liquid leaving the coupling, and does not of itself introduce a loss. Russian Scoop- Tube Coupling. An interesting variation of the scoop tube principle of regulating a variable-filling coupling is illustrated by Fig. 11, showing a design of Russian origin recently published.\" The shell of the runner is omitted, with the object of reducing the hydraulic thrust, and this offers advantages with the bearing arrangements shown in the drawing. On the other hand, during manufacture and erection of such a coupling it is rather easy for the runner to be dropped and for one or other of the radial projecting vanes to be cracked, with the risk of subsequent fracture when running under load. The vanes of the conventional design of runner are effectively protected by the shell and no problem is presented in dealing with the hydraulic thrust by such simple means as a duplex bearing with tie rod as shown in Fig. 19. The most remarkable feature of the design in Fig. 11 is that the scoop tube is * Soviet Boiler and Turbine Building, 1937, No. 7. 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 95 located inside the working circuit near the inner profile diameter, and carries a swivel head with a \u201cweather vane\u201d to ensure that it presents an open mouth to any liquid circulating in its neighbourhood. It appears that the coupling is designed to transmit 870 h.p. at 730 r.p.m. for driving an induced draught fan in a large boiler house",
            " Experience shows that fluid couplings used for this kind of drive usually operate between 60 and 80 per cent of the motor speed when the boiler is steaming at its maximum continuous rating, the reserve of speed being used only to give the overload evaporation under adverse conditions, such as when the boiler is dirty and the carbon dioxide percentage low. On the other hand, when steaming at light load the couplings are required to regulate down to about 20 per cent of the motor speed. In the coupling shown in Fig. 11, however, the scoop tube is situated in a 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from region where no liquid would be picked up by such a scoop under the ordinary fan drive conditions described, and hence it does not appear possible to give any useful range of speed regulation with this design. The reliability of a main power transmission is of course determined by the reliability of any auxiliaries essential to its operation. In the case of the pump-controlled scoop tube coupling shown by Fig",
            " The use of cooling water was, however, exaggerated by advocates of other methods of draught control and it was not widely appreciated that the quantity required was small, and was frequently available from an existing service, e.g. the cooling water for the fan bearings. Mr. Sykes had raised several interesting questions. As regards thrust equalizing ports, he did not favour the elimination of the shell of the runner for several reasons. In the first place it was desirable to have a positive axial location bearing, and the tie-rod type of thrust bearing was much superior to the arrangement shown in Fig. 11, p. 95, as it was easily able to deal with the normal hydraulic thrust. If specially desired, the thrust could be reduced by the use of equalizing ports through the shell, without weakening the runner materially. On the 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 181 other hand, the elimination of the shell involved the disadvantage that fine running clearances had to be maintained between the unshrouded vanes and the casing, otherwise the slip became excessive ; furthermore, the vanes without the runner shell projected radially like spokes and were no longer protected, as in the standard design, from accidental damage which might be overlooked by an inspector. So far as he was aware the design referred to by Mr. Sykes had been developed without outside assistance and in his opinion it possessed nothing but disadvantages compared with the standard designs, which were based upon the collaboration and pooled experience of an international group. The disadvantage of the scoop arrangement shown in Fig. 11 was insurmountable for fan drives, since a scoop tube in the position shown could not possibly transfer liquid from the working circuit to effect any useful range of speed control. The diaphragm type of quick-emptying valve shown in Fig. 20 was closed by admitting liquid on the top or radially outermost side of the diaphragm, which was of larger area than the under side exposed to the pressure of liquid in the working circuit. T o open the diaphragm valve the supply of liquid to the top side was shut off, with the result that the small quantity of liquid in the filling passage leaked off centrifugally through small ports and the internal pressure overcame the control pressure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.17-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.17-1.png",
        "caption": "Fig. 10.17 Process flow for continuous production of NC",
        "texts": [
            " Actual yield = The mass of final nitrocellulose/The mass of refined cellulose 100%. Since the chemical reactions and mechanical transportation process in the production of NC cause a certain loss of product, the actual yield in production is always less than the theoretical yield. The yield in production is an important indicator of the rationality of the production process. 5. Nitrocellulose production process [1\u20133, 7] The process of producing NC with nitric and sulfuric mixed acid is shown in Fig. 10.17. Figure 10.17 is the layout of U-type continuous nitration production line for NC, which are composed of 22 sections, including disintegrator, air drying pipe, flowmeter, cyclone separator, U-type nitrator, horizontal acid removal machine, buffering washer, NC pump, digestion barrels, transfer tank, water filter, shredder, fine washer, blender, centricleaner, concentrator, exchange tank, water removal machine, packaging metric equipment, and so on. After nitrating mixed acid is ejected from the sprayer to impregnate the fiber material, mixed acid and cellulose move forward along the U-shaped tube and are discharged from the last U-tube into the horizontal piston discharge centrifuge"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003766_0094-114x(95)00069-b-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003766_0094-114x(95)00069-b-Figure3-1.png",
        "caption": "Fig. 3. Change of the lever arm between a joint and the base frame {bs} or the end effector frame {ee}. Rotation of a joint j closer to the base than joint i changes the lever arm between joint i and the base. On the other hand, rotating i does not change the lever arm between joint j and the base. Rotation of a joint j closer to the base than joint i does not change the lever arm between joint i and the end effector frame {ee}, as seen from {ee}. On the other hand, rotating i does change the lever arm between joint j and {ee }.",
        "texts": [
            " (16) Serial kinematic chains velocity mapping 141 The derivatives with respect to time or with respect to a joint angle j make use o f the lever arm between j o i n t j and the velocity reference point on the end effector. (Remember that, in the inertial representation, this velocity reference point coincides with the origin o f the base frame!) Rota t ing a joint j that is \" fa r ther\" away f rom the base o f the kinematic chain than joint i does not change the lever a n n between base and joint i, neither in magni tude nor in direction, Fig. 3. Hence, by the definitions (10) and (11), the derivative ~J~/~qJ vanishes in the inertial representation, but not in the body-fixed or hybrid representations. A similar reasoning applies to a joint that is \"closer\" to the base, Fig. 3: body-fixed derivatives vanish, inertial derivatives don ' t . In the \"hybr id\" representation, i.e., bsJ \u0300e, the lever a n n between base and end effector changes, in general, for all combinat ions o f i and j. Table 2 summarizes these properties. 142 H. B r u y n i n c k x a n d J. de S c h u t t e r 4.2. Partial derivatives for R Equations (15) and (16) applied to the rotation matrix between {ee} and {bs} gives 0~R = [b~e j x ] ~eR, 8q y ee bs since only motion of joint j is to be considered"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001752_j.mechmachtheory.2020.103992-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001752_j.mechmachtheory.2020.103992-Figure5-1.png",
        "caption": "Fig. 5. Orthogonal coordinate system on (1) .",
        "texts": [
            " Therefore, in this section, from the point of tensor analysis, the curvature relationships between tooth surfaces are reargued. 3.1. Curvatures and torsion of (1) along (1) From Eqs. (7) \u2013(9) , it can be concluded that the first-order partial derivative of S (1) with respect to u , S (1) u , is another tangent vector of (1) and is orthogonal to r (1) \u2032 along (1) . Let S (1) u = n \u00d7 r (1) \u2032 along (1) . The curvatures of (1) along (1) can then be derived from the variations of the orthogonal coordinate systems constructed by r (1) \u2032 , n \u00d7 r (1) \u2032 , and n as shown in Fig. 5 . According to the position vector representation of (1) in Eqs. (7) \u2013(9) , along (1) , the first-order partial derivatives of S (1) are obtained as S (1) t = \u2202S (1) (t,u ) \u2202t \u2223\u2223\u2223 u =0 = d S (1) (t, 0) d t = r (1) \u2032 S (1) u = \u2202S (1) (t,u ) \u2202u \u2223\u2223\u2223 u =0 = n \u00d7 r (1) \u2032 , (32) where k (1) is a variable coefficient. Therefore, the contravariant components of the metric tensor of (1) along (1) can be obtained as a (1) 11 = S (1) t S (1) t = r (1) \u2032 r (1) \u2032 a (1) 12 = a (1) 21 = S (1) t S (1) u = r (1) \u2032 (n \u00d7 r (1) \u2032 ) = 0 a (1) 22 = S (1) u S (1) u = (n \u00d7 r (1) \u2032 ) (n \u00d7 r (1) \u2032 ) = (n n )( r (1) \u2032 r (1) \u2032 ) \u2212 (n r (1) \u2032 )(n r (1) \u2032 ) = r (1) \u2032 r (1) \u2032 = a (1) 11 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002458_j.promfg.2020.08.016-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002458_j.promfg.2020.08.016-Figure9-1.png",
        "caption": "Fig. 9. Numerical thermal field during the deposition of the layer number 16; black area corresponds to the melt pool, and location of 3 points of interests.",
        "texts": [
            " This result confirms the small effect of a slight variation of conductivity and it is justified by the high conductivity of the clad, reached by the addition of 20% of tungsten carbides to the 316L matrix. Similarly, a variation of 3.5% in the amount of the apparent heat capacity of the clad (possible measurement error) is shown in Fig. 8. It results in a maximum difference of 16.5\u00b0C. This impact of a slight perturbation could be due to two different reasons; the high conduction of the clad on the one hand, and the high volume of the substrate (compared to the clad) on the other hand. The computed thermal field during the deposition of the 16th layer is illustrated in Fig. 9. In Fig. 10, we propose to follow the thermal history of three points of interest representing two adjacent elements at each time: in the beginning of deposition (point 1), in the medium (point 2) and towards the end of the deposition (point 3). When the laser arrives to the midpoint of the 4th layer, the temperature is about 1730\u00b0C, this value slightly increases when reaching midpoints of layers 10 and 16 (about 1800 \u00b0C). These values exceed the melting point of the 316L matrix (1400\u00b0C), but are below the melting point of WC carbides (2800\u00b0C)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000488_elan.201501095-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000488_elan.201501095-Figure5-1.png",
        "caption": "Fig. 5. DPV curves recorded for the GCE/Naf/CuO in 0.01 mM carbofuran (a) against 100-fold of common interferents (b) number of repetitive runs and (c) for the stability during three months of storage period under normal room temperature conditions.",
        "texts": [
            " and LOQ were estimated as three and ten times the standard deviation (3s/slope)(10s/slope) of blank DPV signal respectively. The selectivity of the developed sensor was evaluated by recording its response in 0.01 mM carbofuran solution in presence of common co-existing interferents classified as cations, anions and other carbamates. The interferents were pre-treated in 0.1 M NaOH at 40 8C to mimic sensing environment as mentioned in section 3.4. The variation in the measured current response of GCE/Naf/CuO against 100-fold of interferents can be seen in recorded DPV curves shown in Figure 5(a). Although, the developed sensor system demonstrated excellent anti-interference capability for various cation and anions, slight oxidation of other carbamates was observed around higher potential values. However, the observed voltage difference in between the carbamates and the intensity of the generated response was large enough to selectively quantify carbofuran in a complex matrix system. The reliability of the sensor system was evaluated by taking repetitive runs in 0.01 mM carbofuran solution. Figure 5(b) shows the normalised current response for GCE/ Naf/CuO against number of repetitive runs. The estimated RSD values of <1.0 % during 10 repetitive measurements signify the excellent reproducibility of the developed sensor. The stability was evaluated by measuring the current response of GCE/Naf/CuO for 0.01 mM carbofuran within the selected buffer solution during the 3 months of storage period under normal room temperature conditions. The generated bar graphs shown in Figure 5(c) with RSD value (<1.2%) provide evidence of the excellent stability associated with the developed sensor. The performance capability of the developed sensor system for practical application was evaluated by monitoring its potential to quantify carbofuran in various vegetables extracts. Since the collected samples were determined to be either free or contaminated with lower levels of carbofuran, they were found insensitive to the developed sensor. Henceforth, samples were spiked with certain amount of standard carbofuran which was later determined using the developed assay"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002328_s13369-020-04563-x-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002328_s13369-020-04563-x-Figure1-1.png",
        "caption": "Fig. 1 Prosthetic leg model with rigid ankle [2]",
        "texts": [
            " 2, the prosthetic leg model is discussed. In Sect. 3, the SDRE and sliding mode controller and their formulations are addressed. Section 4 presents the proposed method on the robot/prosthesis system. The simulation results for both nominal and uncertain systems are discussed in Sect. 5. Finally, Sect. 6 concludes the paper. This section represents the proposed model for the threerigid links prosthetic leg with three DOF. The model has a Prismatic-Revolute-Revolute (PRR) structure. As shown in Fig. 1, the degree of vertical freedom represents the human hip motion, first and second rotational axis motions representing human angular thigh motion and the prosthetic knee angular motion, respectively. The three-DOF model can be written as (1): (1) M(q(t))q\u0308(t) + Cp(q(t), q\u0307(t))q\u0307(t) + Gp(q(t)) + Rp(q(t)) u(t) \u2212 Te where qT [ q1 q2 q3 ] is the generalized displacement vector of the joint (q1 is the hip vertical displacement, q2 is the thigh angle, and q3 is the knee angle). M(q) is the inertial matrix, Cp(q \u00b7 q\u0307) is the Coriolis and centripetal matrix, Gp (q) is the gravity vector, and Rp(q) is the nonlinear dumping vector; u is the control signal that consists of the active control force at the hip and the active control torques at the thigh and knee; Te is the combined effect of the horizontal component Fx and vertical component Fz of the GRF on each joint",
            " (2), (3), (4), and (5): Te \u23a1 \u23a3 Fz Fz(l2 cos(q2) + l3 cos(q2 + q3)) \u2212 Fx (l2 sin(q2) + l3 sin(q2 + q3)) Fz(l3 cos(q2 + q3)) \u2212 Fx (l3 sin(q2 + q3)) \u23a4 \u23a6 (2) Lz q1 + l2 sin(q2) + l3 sin(q2 + q3) (3) Fz { 0, Lz < sz kb(sz \u2212 Lz), Lz > sz (4) where Lz denotes the vertical position from the lower leg in the overall frame (x0,y0,z0), l2 is the thigh length, and l3 is shank length; Sz is the vertical distance between the base of the frame and the belt (treadmill standoff); kb is the belt stiffness; and \u03b2 is the friction coefficient of the belt (Fig. 1). The nominal values of model parameters and the nominal system parameter values are given in Tables 1 and 2, respectively. Finally, the states, control input, and desired trajectories are presented in Eq. (6) as follows: xT [ q1 q2 q3 q\u03071 q\u03072 q\u03073 ] uT [ fhip \u03c4thigh \u03c4knee ] rT [ rd1 rd2 rd3 rd4 rd5 rd6 ] [ qd2 qd2 qd3 q\u0307d1 q\u0307d2 q\u0307d3 ] (6) Next, it is attempted to design a controller that minimizes power consumption and the state variables q1, q2, and q3 follow a predetermined path with high accuracy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001207_rpj-07-2018-0171-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001207_rpj-07-2018-0171-Figure4-1.png",
        "caption": "Figure 4 Optimization for additive manufacturing",
        "texts": [
            " In this design stage, the supporting beam was added to prevent a flange fromdeformation [Figure 3(c)]. The basic shape of the axle carrier was then upgraded during three follow-up design stages. The component orientation on the building platform was chosen and the shape was additionally optimized with regard to additive manufacturing Topologically optimized axle carrier Ond rej Vaverka, Daniel Koutny and David Palousek Rapid Prototyping Journal Volume 25 \u00b7 Number 9 \u00b7 2019 \u00b7 1545\u20131551 and expected machining. Areas that should be milled (under bolts) were elevated for easier machining [Figure 4(a)]. Cylindrical parts were angled towards the building platform (drop-like shape was obtained), so the need for supports was reduced [Figure 4(b)]. Chamfers were added instead of rounded edges [Figure 4(c)]. Long supporting parts were angled towards the building platform if the loading direction allowed it [Figure 4(d)]. The development of the objective function with all intermediate designs in every design stage is introduced in Figure 5. The axle carrier model was simplified for the FEM simulations (reducing chamfers and small radii) for improving the calculation stability and mesh quality. The adaptive meshing method with tetrahedral elements of size 4mm was used. Components that were not the aim of the analysis were significantly simplified and the elements of sizes 3mm (bolts and bearings) and 10mm (brake calliper) were prescribed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002572_j.apm.2020.12.020-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002572_j.apm.2020.12.020-Figure6-1.png",
        "caption": "Fig. 6. First three orders of vibratory mode under Pose I.",
        "texts": [
            " Since the number of flexible beam elements in ANSYS software is far more than that used in the dynamic model of this paper, the simulation deviations are acceptable. The vibratory mode schematic diagrams of the Table 2 Frequencies under Pose I (Unit: Hz). Order No. First order Second order Third order Theoretical elasto-dynamics model 9.279 9.282 29.622 ANSYS software simulation 9.829 10.138 33.332 Table 3 Frequencies under Pose II (Unit: Hz). Order No. First order Second order Third order Theoretical elasto-dynamics model 7.188 10.352 28.595 ANSYS software simulation 7.866 9.342 31.622 first three natural frequencies are shown in Fig. 6 to Fig. 7 , in which the first order of vibratory mode is the mechanism vibrating left and right along the y axis, the second order of vibratory mode is the mechanism vibrating back and forth along the x axis, and the third order of vibratory mode is the torsional vibration of the mechanism around the z axis. In general, the Lagrangian multiplier must be solved simultaneously when solving the system dynamic equations of index-1, which inevitably increases the complexity of the solving process. In this section, we propose a dynamic equation solving strategy which combines Baumgarte stabilization technique and Gill algorithm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure7-1.png",
        "caption": "Fig. 7 3D model of slicing cutter with plane rake face",
        "texts": [
            " Based on the above parameters, the cutting edge, the curved rake face, and the flank face are calculated by the method described in Sections 2.2 and 2.3. And the 3D model is generated as shown in Fig. 6. Different from design angles (or called shape angles), the working angles are dynamic in regard to the cutting velocity, which directly affects the cutting process. The working angles defined in the orthogonal plane can better reflect the cutting performance of the cutter. Taking the workpiece in Section 2.4 as an example, the cutter with a plane rake face is also designed as shown in Fig. 7, and the comparison with the curved rake face is shown in Fig. 8. For convenience, the cutter with curved rake face is named cutter A, and the cutter with plane rake face is named cutter B. It is obvious that the working rake angle of cutter A is just equal to \u03b3o since the rake face takes the working rake angle as the design variable directly. However, the working rake angle of cutter B is not distinct since its rake face is designed according to the shape rake angle \u03b3e. For comparison, the working rake angle of cutter B is calculated by use of the reference system as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002412_0142331220941934-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002412_0142331220941934-Figure1-1.png",
        "caption": "Figure 1. Relative motion geometry between target and vehicle.",
        "texts": [
            " The global stability and all subsystems stability of the FTIGC system are proven by Lyapunov stability theory. The rest of this paper is organized as follows. First, the IGC model derivation is described. The next section formulates the FTIGC scheme and analyses its stability, followed by verification of the effectiveness and robustness of the FTIGC system through simulations. The final section provides the conclusions. Mathematical model 3D relative dynamics The 3D relative dynamics of hypersonic vehicles and the target are established. The relative motion geometry is provided in Figure 1. Oxyz is the inertial coordinate system, and Oxsyszs is the line-of-sight (LOS) frame system. O represents the centroid of the hypersonic vehicle, T is the centroid of the target, R is the distance of O and T , qe is the angle of elevation of the LOS, and qh is the angle of the azimuth of the LOS. Wang et al. (2016b) described these frame systems. The relative dynamic equation (Wang et al., 2019) is expressed as follows: \u20acR=R _q2 h cos q2 e +R _q2 e + aTr aMr \u20acqe = 2 _R _qe R _q2 h cos qe sin qe + aTe aMe R \u20acqh = 2 _R _qh R + 2 _qe _qh tan qe aTh aMh R cos qh 8>< >>: , \u00f01\u00de where aTr, aTe, aTh are the components of the target acceleration along the x-, y- and z-axes in the LOS frame system, and aMr, aMe, aMh are the components of hypersonic vehicle acceleration along the x-, y- and z-axes in the LOS frame system, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000325_3387304.3387327-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000325_3387304.3387327-Figure1-1.png",
        "caption": "Figure 1. Quadcopter configuration.",
        "texts": [
            " To verify the proposed algorithm, several numerical and hardwarein-the-loop simulations are carried out. A DJI-F450 multicopter is used as the experimental platform to conduct on-testbed attitude flight for the evaluation of the effectiveness of our method. ORMULATION 2.1 Quadcopter Dynamics Model Here, we briefly summarize the quadcopter attitude dynamics model as it has already been presented previously [1], [7]-[9]. Let ,  , and  denote the three Euler angles roll, pitch, and yaw, respectively, where, / 2  , / 2  , and   (see Fig. 1). ,x yJ J , and zJ the moments of inertia along the ,x y , and z axes, respectively; l the arm length of the vehicle. The quadcopter dynamics model can be described as follows: Permission to make digital or hard copies of all or part of this wor k for personal or classroom use is granted without fee provided that copies are not made or distributed for profit or commercial advantag e and that copies bear this notice and the full citation on the first page. Copyrights for components of this work owned by others than ACM must be honored"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001798_s00419-020-01733-z-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001798_s00419-020-01733-z-Figure4-1.png",
        "caption": "Fig. 4 Section view of defect geometry in the outer race",
        "texts": [
            "When the instantaneous contact point of the ball is outside the defect area, Rv = R. But if the ball is located in the defect zone, Rv is to be calculated. Using trigonometry equation, the following three equations can be written: c = \u221a R2 v+(R + t)2 \u2212 2Rv(R + t) cos \u03b8 \u2032, c = oa (37) c = \u221a r2t +r 2 t \u2212 2r2t cos \u03b8 \u2032\u2032 (38) Rv = \u221a (R \u2212 rt + t)2+r2t \u2212 2rt(R \u2212 rt + t) cos \u03b8 \u2032\u2032 (39) With simultaneous solving of the above three equations, the variable radius of the outer race (Rv) at different angles shown in Fig. 4 can be calculated. To verify the model, experimental tests have been set in different rotating speeds. The experimental test in cases of healthy bearing, and bearing with spall defect on the outer race are performed. For this purpose, a test bench was designed with the following characteristics: 1. Using a small electric motor with a maximum speed of 2825 rpm and 0.37 kW 2. Optimized design and manufacturing of precision bearing housing 3. Using the angular contact ball bearings SKF 7202-BEP Angular contact ball bearings due to their specific design should not be used alone"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003954_004051750107100113-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003954_004051750107100113-Figure2-1.png",
        "caption": "FIGURE 2. Principle of fiber migration measurement [16].",
        "texts": [
            " The principle of measuring and analyzing the fiber trajectory is based on the tracer fiber method suggested by Riding [16]. On immersion of the yarn into a liquid with the same refractive index as that of the specimen, the white fibers composing the yarn become almost optically dissolved and the colored tracer fibers can be readily observed. A mirror, placed in the liquid vessel at 45 degrees to the direction of the incident light, which is arranged perpendicular to the observation camera, simultaneously creates two perpendicular cross-sectional images for the yarn (Figure 2). The immersion liquid used for the measurement is methyl salicylate. After the tracer is observed on the monitor with the CCD camera, images for the whole length of each tracer fiber in yarn are captured by the vision control system, and the the the image data are stored in the computer. Figure 3 shows typical images of a tracer fiber in the yam, obtained during measurement. at Bobst Library, New York University on May 7, 2015trj.sagepub.comDownloaded from re-established to transform the image data in such a way as to yield the trajectory of the tracer fibers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000046_j.optlastec.2019.105728-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000046_j.optlastec.2019.105728-Figure1-1.png",
        "caption": "Fig. 1. Laser-CMT hybrid welding setup, and sketch of the geometrical parameters.",
        "texts": [
            " The specimen surfaces were chemically cleaned by acetone before welding to eliminate surface contamination. The NiCu 1-IG filler wire with 1.2 mm in diameter was used. The chemical compositions of the base metal and the filler wire were shown in Table 1. The laser-CMT hybrid welding system include a 3 kW CO2 laser source and Fronius CMT-2700 welding machine. During the experiment, the CMT welding torch and the laser head are fixed together with an angle about 45\u00b0 and the laser beam was placed before the CMT arc (Fig. 1). Because the base metal was low carbon steel and the transfer form was short circuit transfer, so the shielding gas was 80% volume argon (Ar) mix with 20% volume CO2. To evaluate the weld experiments, high speed imaging at 5000 frames/s was also used during the welding process for each experiment to observe the melt flow and droplet transfer behavior. The transfer frequency of the droplet was calculated by Image-Pro Plus image processing software with more than 10,000 frames, and the brightness curve of the arc area was obtained. Significant changes of the droplet volume and shape were observed with different parameters. Therefore, the projection area measurement of the droplet was used as a substitute for the calculation of the droplet volume, as shown in Fig. 1(c). After welding experiment, the samples were processed by wire cutting. It has been grinded with sandpaper for many times and polished. Finally, 4% nitric acid alcohol was used for corrosion, and VHX1000E optical microscope was used for observation. The laser plasma spectroscopy was measured by Avantes 2048 spectrometer. Then the Boltzmann diagram method is used to calculate the laser plasma temperature. First, the characteristics of CMT droplet transfer without the laser were studied via an orthogonal experiment (as shown in table 2)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure22-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure22-1.png",
        "caption": "Fig. 22. Two axial pistons used in the bottom side of the reference pump design to balance the axial loads. For each piston, the inner surface is in contact with the shaft end on the bottom side, and the outer surface of each piston is connected to the delivery discharge chamber with high pressure.",
        "texts": [
            " Next, the second scenario \u201caxial dynamic\u201d is studied to design the proper balancing mechanism for the reference pump. Under this scenario, without a properly designed axial supporting mechanism, the gears will be pushed to the bottom side and wear out the bottom bushings quickly, leaving a large axial gap on the top side at the same time. Such a situation in practice will result in very low mechanical efficiency, volumetric efficiency, as well as a short lifetime of the machinery. The studied axial supporting mechanism is to use two axial pistons, one for each gear, on the bottom side (see Fig. 22). The inner X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 face of each piston is in contact with the bottom-end surface of the gear shaft, and the outer side is connected to the high-pressure discharge chamber on the delivery side. This is a hydrostatic balancing method such that the supporting force will adjust itself according to the operating condition. The pistons are assumed to be two cylinders, and the diameters of the cylinders Dap need to be designed properly to provide the supporting forces as needed",
            " Based on this model, design optimization can be made. X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 This subsection presents the experimental evidence for the validity of the simulation model discussed throughout this paper. The experimental measurement was conducted on a commercial CCHGP unit produced by Settima Meccanica. The pump under test was a 7-tooth gear pump with a 22 cm3/rev displacement, shown in Fig. 24. The implementation of the axial balancing mechanism is very similar to the design shown in Fig. 22, discussed in the preceding subsection. The tests were performed at the authors\u2019 research center of Purdue University. The fluid used during the test is an ISO VG 46 hydraulic oil. All measurements were performed at 50 \u25e6C. The detailed information about the test setup, including sensors, hydraulic circuits, etc., is described in detail in the authors\u2019 previous paper [20], which is dedicated to discussing the fluid dynamic behavior of CCHGP pump design. However, predicting the micro-motion of the solid parts is a necessary step in accurately modeling the efficiencies of a hydraulic pump"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001535_j.ymssp.2020.106723-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001535_j.ymssp.2020.106723-Figure16-1.png",
        "caption": "Fig. 16. Schematic of experimental universal joint.",
        "texts": [
            " To mimic the transmission system of CRH train, the distance between centers of universal joints is designed at 676 mm and the height is designed at 52 mm, 5. Drive shaft, 6. AC motor, 7. Force sensor, and 8. Signal acquisition unit. Fig. 15. Experimental set-up. as such the articulation angle of the designed cardan shaft works out to be 4.4 degrees, which is the same articulation angle of cardan shaft experienced in CRH train. To simulate cardan shaft misalignment, the depth of step hole is adjustable in movable plates with different designs as shown in Fig. 16. The depth of step hole of mounting cross-shaft in the upper plate is designed shallower than the step hole in the lower plate (see Fig. 16). In this way, the center of cross-shaft is shifted down with a certain distance and the structure of manufactured cardan shaft is like the actual misaligned cardan shaft. The simulated offsets in the experiment are listed in Table 7. Considering the influence of rotational speed on the generated inertial force and the effect of the index, three different rotational speeds are applied in experiment, where the operating speed and other basic parameters of the experimental cardan shaft are listed in Table 8"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001746_j.procs.2020.05.068-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001746_j.procs.2020.05.068-Figure4-1.png",
        "caption": "Fig. 4. Modeling of product",
        "texts": [
            " Based on need student-generated 3 to 4 solutions like pressing system with hydraulics and pneumatics, single two rollers rolling system with an adjustable cutter. The by applying the criteria of compact, cost-effective, production rate, user-friendly and quality of product the students have selected the best solution of two rolling systems with a cutter having an adjustment to change the size and shape of products. Then by considering the constraints and with some initial assumptions students designed the system and modeled with the help of software shown in Figure 4. The product is developed then as per the design by using the facilities provided in the college campus, in addition, the product design is checked from industrial expert to reduce cost and easy for manufacturing and assembly, the suggestions are implemented in the development process for betterment of the product. The final developed product is tested for given constraints and need, then demonstrated in front of customers to ratify that the customer need is satisfied. The outcomes of the course are assessed by the evaluation rubrics designed for each phase of the design-build experience"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003804_(sici)1521-4109(199809)10:12<803::aid-elan803>3.0.co;2-3-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003804_(sici)1521-4109(199809)10:12<803::aid-elan803>3.0.co;2-3-Figure1-1.png",
        "caption": "Fig. 1. Design of the thick film electrodes used as base electrodes for the enzyme printing.",
        "texts": [
            "2 % PEI and 16 mg of graphite powder. This mixture was dried under the same conditions as described above and then mixed with different amounts of the printing polymer (ratio 1:2 and 1:4). In a variation, the same procedure was used but without polyethyleneimine. As base electrodes for the printing process, prefabricated thick film electrodes from GBF (Braunschweig, Germany) were used. The electrodes were fabricated according to the process described by Rohm et al. [20]. The layout is shown in Figure 1. Compared to previous publications the working electrodes were printed using a carbon paste [C-458(J), Ercon, Waltham, MA, USA] which was dried at 150 8C. The enzyme/graphite/polymer mixture was printed on these working electrodes and hardened by exposure to UV irradiation with a mercury vapor lamp for 3\u20135 s. The modification was applied only to one of the carbon electrodes which was used as working electrode. The printed electrodes were then stored dry at 4 8C until required. Some electrodes were additionally covered by placing 2 mL of a polyurethane solution onto the modified electrode and left to dry for about 1 h at room temperature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001207_rpj-07-2018-0171-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001207_rpj-07-2018-0171-Figure2-1.png",
        "caption": "Figure 2 (a) Design space for preliminary study; (b) wrong material distribution; (c) good material distribution",
        "texts": [
            " Optimization results representing optimal component shape were exported in the stereolithography (STL) format. Catia V5R20 (Dassault Syst\u00e8mes, V\u00e9lizy-Villacoublay Cedex, France) was used for the creation of a solid CAD model from the polygonal mesh. Using surface modelling, the required shape was obtained. The FEM analysis was carried out using ANSYS Workbench 17.2 (ANSYS, Inc., Canonsburg, PA, USA).Model from previous design iteration was then used as a design space for next optimization. During a preliminary study on the previous version of the axle carrier [see the design space in Figure 2(a)], it was found out that more realistic results were obtained if the reaction forces were applied in the suspension pickup points and the constraints were applied on the bearings. Otherwise, the optimization algorithm did not connect the point for mounting the steering rod [Figure 2(b)], because constraints A and B (Figure 1) were sufficient for preventing the model from move and constraint C (Figure 1) was in that case redundant. If the forces were applied in the suspension pickup points, this phenomenon did not occur [Figure 2(c)]. Basically, the material was more functionally distributed near the forces than constraints. The finer material distribution was also obtained if the optimization was carried out in iterations and the material was removed gradually. Therefore three design stages were carried out using topology optimization until the weight was less than 500 g (Figure 5). After that, two minor design stages based on FEM evaluation were performed, and the shape of the axle carrier was optimized for additive manufacturing and defects in design were repaired"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure3.1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure3.1-1.png",
        "caption": "Fig. 3.1 Structure of WLLS [8]",
        "texts": [
            " The simulation results verify that the WLLS has improved dynamic response speed and accuracy after adding a controller. The remainder of this paper is organized as follows: the composition, the motion principle, and the mathematical description of WLLS are shown in Sect. 3.2. In Sect. 3.3, the design of sliding mode controller is presented. The fuzzy rules for the parameters are in sliding mode controller given in Sect. 3.4. Section 3.5 shows the simulation results of WLLS under the fuzzy sliding mode controller, and the conclusions are drawn in Sect. 3.6. The general structure of the WLLS is shown in Fig. 3.1 [8]. TheWLLS includes two segments: a vibrating ring-like segment and a contact ring-like segment. The longitudinal springs connect two segments and each segment consists of two ferromagnetic semi-rings. In contact segment, the transverse springs connect two semi-rings. There is also a supporting pad in this part to adapt the material of the pipe wall. The exciting coils, which connect with a power supply, are winded on the semi-rings of the vibrating segment [8, 14, 15]. This design can be simplified so that a magnetic system of the WLLS can be very similar to clapping electromagnet [14, 16]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001184_j.ijfatigue.2019.105281-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001184_j.ijfatigue.2019.105281-Figure4-1.png",
        "caption": "Fig. 4. DC specimen used for the mode II crack growth test. (a) Shape and dimensions of the specimen and (b) the cut direction of the specimen.",
        "texts": [
            " Analysis of the fracture surface: The fracture surface of the specimen was observed with a stereomicroscope and SEM, and the shapes and sizes of the macroscopic defects were analysed. The EDX was used to analyse the chemical compositions of the macroscopic defects. The DC specimen designed by Murakami and Hamada was used to measure the \u0394KII,th of CL60 wheel steel [21,22]. Fig. 3 shows the basic principle of the mode II crack growth test. The direct stress in the neutral section is zero, while the shear stress reaches the maximum in this section. Thereby, a mode II fatigue crack would propagate along the neutral section. Fig. 4 shows the shape and dimensions of the specimen, which was machined from the wheel rim, with the neutral section parallel to the wheel tread. The specimen has a chevron notch (detail A shown in Fig. 4a) and side grooves (detail B shown in Fig. 4a). The fatigue crack is initiated at the tip of the chevron notch. The 60\u00b0 V-shaped side groove is designed to cause mode II fatigue crack growth along the neutral section. For a constant amplitude load, \u0394KII of the crack along the neutral section gradually decreases with the growth of fatigue crack. The mode II threshold stress intensity factor range \u0394KII,th can be calculated if the length of the arrested crack is measured in the test. The setup for the test is shown in Fig. 5. A pair of specimens was tested at the same time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002779_s42243-020-00528-4-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002779_s42243-020-00528-4-Figure3-1.png",
        "caption": "Fig. 3 Stress diagram of strip. a Forward slide zone; b backward slide zone. rx Normal stress",
        "texts": [
            " u\u00f0x\u00dey\u00f0x\u00de \u00bc ueye 2 xe x\u00f0 \u00de _yc \u00fe ye y\u00f0x\u00de\u00f0 \u00de _xc \u00f013\u00de u\u00f0x\u00de \u00bc ueye 2 xe xn\u00f0 \u00de _yc \u00fe ye y\u00f0x\u00de\u00f0 \u00de _xc y\u00f0x\u00de \u00bc vR \u00f014\u00de The neutral point coordinates can be obtained from Eq. (14). xn \u00bc xc \u00fe _ycR vR \u00fe _xc \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi CR vR \u00fe _xc \u00fe _ycR vR \u00fe _xc 2 s \u00f015\u00de where C \u00bc ueye 2\u00f0xe xc\u00de _yc \u00fe ye y0d _xc vRy 0 d and y0d \u00bc yd \u00fe 2yc. According to the slab analysis method, an arbitrary microelement in the deformation area is shown in Fig. 3. According to the Karman\u2019s theory of force balance, a differential equation of p and y(x) can be obtained p\u00fe r\u00f0 \u00de dy\u00f0x\u00de dx \u00fe y\u00f0x\u00de dr dx 2ss \u00bc 0 \u00f016\u00de when x\\ xn, take the negative and when xn\\ x, take the positive. Substituting Eq. (7) into Eq. (16) and simplifying it, the following equations could be obtained r dy\u00f0x\u00de dx y\u00f0x\u00de dp dx 2ss \u00bc 0 \u00f017\u00de ss \u00bc ksa \u00fe 1 k\u00f0 \u00desb \u00f018\u00de sa of the rough contact surface boundary can be calcu- lated by adhesive friction theory. sa \u00bc r 2 \u00f019\u00de sb can be calculated according to Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003678_a:1008896010368-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003678_a:1008896010368-Figure1-1.png",
        "caption": "Figure 1. Biped robot during the double support phase.",
        "texts": [
            " However, to stabilize the biped robot over a wide range of walking speeds (including a \u201czero-speed\u201d stationary phase), the double support phase must be taken into consideration. Motion of a biped robot in the double support phase can be described as the motion of a robot under holonomic constraints if the contact positions between the feet and ground are fixed in the world frame. This condition can be expressed as a set of holonomic constraints (Hemami and Wyman, 1979; Peng and Adachi, 1992). Figure 1 shows the model of our biped robot during the double support phase. The following assumptions are made for the sake of simplicity: A1. Each foot is in contact with the ground along its entire lower surface. A2. The pitch angle of the trunk, q5, is fixed during walking. Assumption A1 implies that it is possible to supply an arbitrary input torque to the robot through the ankle joint. Assumption A2 means that the inertia force due to the rotation of the trunk is negligible in the dynamic equations of the robot. These assumptions are realized by constructing each leg using a pair of parallel links. The structure and actuation of the leg are shown in Fig. 2. Two D.C. motors attached at the trunk supply input torques \u03c41a and \u03c42a in Fig. 2(a) through a timing belt and reduction gear. Since \u03c41a = \u03c41b and \u03c42a = \u03c42b, the leg can be modeled simply as shown in Fig. 2(b). The constraints are expressed as c(q) = 0. (1) Since the right foot in Fig. 1 is fixed on the ground, c(q) = [ l1 sin(q1)+ l2 sin(q2)\u2212 l2 sin(q3)\u2212 l1 sin(q4)\u2212 L l1 cos(q1)+ l2 cos(q2)+ l2 cos(q3)+ l1 cos(q4) ] . (2) The dynamic equation for the robot is written as H(q)q\u0308 + h(q, q\u0307)+ CT (q)\u03bb = \u03c4, q = (q1, q2, q3, q4) T , H \u2208 R4\u00d74, h \u2208 R4\u00d71, \u03bb \u2208 R2\u00d71, \u03c4 \u2208 R4\u00d71, C \u2208 R2\u00d74, (3) P1: VTL/TKL P2: EHE/TKL P3: VTL/TKL QC: PMR/TKJ T1: PMR Autonomous Robots KL465-06-Mitobe May 16, 1997 17:19 Control of a Biped Walking Robot 289 where Hq\u0308 represents the inertial forces, h is the gravitational, centripetal and Coriolis term, \u03bb represents the constraint forces, CT\u03bb represents the torques at each joint due to the constraints, and \u03c4 represents the joint torques",
            " For a system under holonomic constraints, it is possible to introduce a set of independent generalized coordinates to describe the constrained system and to eliminate the constrained forces from the equation of motion (Goldstein, 1980). In the case of our biped robot, the number of degrees of freedom of motion can thus be reduced from 4 to 2. As for motion generation of the biped gait, our control objective is to provide a stable trajectory to the trunk of the robot. In order to control the position of the trunk with respect to the ground, the independent generalized coordinates should be the coordinates of a point fixed on the trunk. The coordinates of a point fixed at the bottom of the trunk (point P in Fig. 1) were selected as the independent generalized coordinates. Since the orientation of the trunk is fixed with respect to the ground (assumption A2), the orientation of the trunk need not be considered in this discussion. The independent generalized coordinates p \u2208 R2\u00d71 can be written as p = d(q) = [ l1 sin(q1)+ l2 sin(q2) l1 cos(q1)+ l2 cos(q2)+ a3 ] . (8) Differentiating Eq. (8) with respect to time yields p\u0307 = D(q)q\u0307, (9) p\u0308 = D(q)q\u0308 + D\u0307(q)q\u0307, (10) where D = \u2202d \u2202q \u2208 R2\u00d74. (11) The matrix D is assumed to be of rank 2 by the assumption that q1\u2212 q2 6= 0 or \u03c0 , and q3\u2212 q4 6= 0 or \u03c0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002799_tie.2021.3063991-Figure17-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002799_tie.2021.3063991-Figure17-1.png",
        "caption": "Fig. 17. Thermal distribution of the proposed machine.",
        "texts": [
            "5 A/mm2 Jf = 0 A/mm2 Jf = -2.5 A/mm2 0 120 240 360 480 600 720 0 1 2 3 4 5 C o re l o ss ( W ) Speed (r/min) Jf = 2.5 A/mm2 Jf = 0 A/mm2 Jf = -2.5 A/mm2 Authorized licensed use limited to: Dedan Kimathi University of Technology. Downloaded on June 28,2021 at 18:41:49 UTC from IEEE Xplore. Restrictions apply. 0278-0046 (c) 2021 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. Fig. 17 shows the thermal distribution of the proposed machine under the rated flux enhancing condition. Namely, the torque, speed, Ja and Jf are 6.43 Nm, 120 r/min, 3.4 A/mm 2 and 2.5 A/mm 2 , respectively. Also, ambient temperature is 20\u00b0C. As can be seen, the maximum temperature 72\u00b0C is located at the armature winding. And, the PM temperature is about 65\u00b0C. In order to check the PM demagnetization, the material property of N35SH at 80\u00b0C is set to the PM. Fig. 18 depicts the flux density distribution in the PMs under the rated flux enhancing condition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002489_j.jwpe.2020.101671-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002489_j.jwpe.2020.101671-Figure8-1.png",
        "caption": "Fig. 8. Reynolds number and calculated shear rate of three different cell models tested in the simulations, elongated without channels or obstacles to substrate flow (1), with 8 channels distributed along with the anodic chamber (2), with 4 channels distributed along with the anodic chamber width (3).",
        "texts": [
            " These numerical simulations facilitate the flow type verification and expose the distribution of the fluids over the channels, to improve this J.D. Lo\u0301pez-Hincapie\u0301 et al. Journal of Water Process Engineering 38 (2020) 101671 characteristic and obtain a desirable Reynolds number [2]. This analysis is carried out focusing on the transport phenomena present in the channel, as the Reynolds number is a quantity that directly intervenes in energy conversion efficiency [14]. Reynolds\u2019s numbers were calculated from the maximum flow velocities obtained to the different architectures of the anode chamber (Fig. 8). The highest value observed was utilized to select the architecture of the MFC. The Reynolds number was calculated using the specific gravity, viscosity, and linear velocity of the fluid, the cross-sectional area, and the wetted perimeter of the anode chamber. For fluid\u2019s specific gravity and viscosity, clean water was assumed. The linear velocity was obtained from the CFD runs while the crosssectional area, and wetted perimeter were calculated considering the internal measurements of the MFCs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001462_iros40897.2019.8967640-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001462_iros40897.2019.8967640-Figure6-1.png",
        "caption": "Fig. 6. Light green triangle (M) is sweeped by member e when moving node v to new location v\u2032 along the yellow (\u2192) trajectory and the new position of member e is e\u2032. e doesn\u2019t collide with any other members in (a) but does collide with two members in (b) during the motion.",
        "texts": [
            " Every edge module can be modeled as a line segment in space, thus, for every e \u2208 Ev , na\u0131\u0308vely we can just check the intersection between a moving line segment and a still line segment. This is actually not easy. With our nonuniform grid space and node motion model, the node is actually moving along a line segment. Hence, we present an easier and more efficient method to do collision check. For a VTT configuration G = (V,E) and a node v \u2208 V , once a reconfiguration action a is executed, every moving member in Ev sweeps a triangle area (Fig. 6a), and if this member e \u2208 Ev collides with another member e\u0304 \u2208 E\\Ev , then e\u0304 must intersect with the triangle M generated by e (Fig. 6b). There are two cases: e\u0304 is not parallel to M and e\u0304 is parallel to M. For the first case, there are already many efficient algorithm to test the intersection between a line segment and a triangle in 3D space, such as Mo\u0308ller-Trumbore ray-triangle intersection algorithm [19]. For the second case, it is just a simple 2D geometry math problem. A transition model is required to describe the effect of a reconfiguration action on a VTT configuration. For a VTT configuration G = (V,E), given a discrete reconfiguration action a, a new VTT configuration G\u2032 = (V \u2032, E\u2032) is able to be computed by transition model F(G, a) if action a is executable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002542_icem49940.2020.9270965-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002542_icem49940.2020.9270965-Figure12-1.png",
        "caption": "Fig. 12. Structure of proposed topology [27].",
        "texts": [
            " These stator-PM machines have additional benefit, which PMs mounted in stator are easier to dissipate heat, especially in post-fault operation. Meanwhile, the hybrid excitation synchronous machines (HESM) for SG application are also get much attention recently [31, 32], in Fig. 11. The common thing in these types is controllable flux to amend in the presence of fault. The fault winding, especially short circuit, of the machine could benefit from the controllable flux. A doubly-salient flux controllable out-rotor motor topology has been presented in [27], Fig. 12. The innovation of this topology is based on double-salient motor and memory motor. Adopted the AlNiCo PMs can be completely demagnetized in the post-fault mode and the motor will operate as switched reluctance motor, which Authorized licensed use limited to: University of Gothenburg. Downloaded on December 20,2020 at 19:19:58 UTC from IEEE Xplore. Restrictions apply. possessed inherently FT capability. Meanwhile, the novel topology has high power density and wide speed range simultaneously 3) Separated Phase Machine An integral slot non-overlapping concentrated winding (ISNCW) and machine configuration have been proposed [28], shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure52-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure52-1.png",
        "caption": "Fig. 52. Comparative Sizes of New Coupling and Vulcan-Sinclair Coupling",
        "texts": [
            " The arrangement shown gave almost unlimited high speeds, and the movable scoop was taken out when it was not needed. It was extremely quick and very strong, and there were no idling losses. While an ordinary centrifugal pump required about 30-35 per cent of its full load at zero, the pump in question, consisting of the tank and the scoop, required nothing. He had built such a gear, and its size was only 50-60 per cent of that of an ordinary coupling. The comparative sizes of the new coupling and the Vulcan-Sinclair gear and other gears were shown in Fig. 52. Dr. J. N. H. TAIT remarked that his experience was limited to the coupling of the constant filling type, generally known as the \u201cfluid flywheel\u201d. The drag torque of that type limited its use for automobile 11 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from work to combination with an epicyclic type of gearbox where drag did not prevent gear engagement, unless an additional controllable friction clutch was provided. With that combination, the fluid flywheel materially improved the handling of an automobile, especially in traffic"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002471_j.engfailanal.2020.104811-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002471_j.engfailanal.2020.104811-Figure11-1.png",
        "caption": "Fig. 11. Modal frequencies and modal shapes of the CDBG without additional constraints arising from gear-teeth meshing.",
        "texts": [
            " This can be formulated as = =f f f f k f k i[ , \u00af ] [ ( ), ( \u00af )] , 1, 2, 3. ..g i g i g i g i c g i c, , , , , (3) where fg i, is the modal frequency of the ith natural mode, and f g i, and f\u0304g i, are the lower and upper boundaries of the interval of fg i, . Based on these boundary conditions, the natural modes and vibration response of the CDBG can be calculated by a commercial FEM solver. When =k 0c , the modal frequencies and modal shapes of the CDBG without the additional constraints of gear-teeth meshing are illustrated in Fig. 11. According to the modal shapes, it can be concluded that, without considering additional constraints, the vibration stress distribution in the gear is axisymmetric about the rotation axis. This evidently does not coincide with the crackpropagation direction in the failed gear. With = \u00d7k N m5 10 /c 9 , the swing and first five modal frequencies of the CDBG attributable to the additional constraints arising from teeth meshing are given in Table 4. For the modes of lower frequencies, the effect of tooth surface meshing significantly increases the natural modal frequency of the CDBG, while for modes of higher frequencies, the effect of the additional constraints is negligible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001443_ls.1500-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001443_ls.1500-Figure2-1.png",
        "caption": "FIGURE 2 Schematic diagram of UMT-3 reciprocating friction and wear tribometer",
        "texts": [
            " Their functional groups were carried out by Fourier transform infrared analysis (FTIR, Nicolet-5700) with the wavenumber range of 400 to 4000 cm\u22121. Thermogravimetric analysis (TGA) was performed by the NETZSCH SAT 449 F3 (NETZSCH Instrument Manufacturing Co, Ltd). The tribological performance was investigated via UMT-3 reciprocating friction and wear tribometer with ball-onplate contact (Center for Tribology Company, USA) at the reciprocating stroke of 2 mm, frequency of 2 Hz, and the applied load from 10 N (Hertzian contact pressure [HCP]: 1.0 GPa) to 60 N (HCP: 1.82 GPa). Figure 2 gives the schematic diagram of UMT-3 tribometer. AISI 52100 steel ball with the diameter of 9.57 mm slides against AISI 52100 steel disc with \u03a6 of 24 \u00d7 7.9 mm. The real-time changes of friction coefficient were automatically recorded by the computer that linked to UMT-3 reciprocating friction and wear tribometer. A 0.5 g of lubricating grease was coated on the contact area of friction pairs. After a few cycles, the grease appeared the \u201cbleeding oil\u201d on the interface of friction pairs under the effect of yield stress"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.37-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.37-1.png",
        "caption": "Fig. 10.37 The structure of cone-shaped chopper [2]: 1, 2\u2014the traditional system, 3\u2014hand wheel regulator, 4\u2014track, 5\u2014shell of chopper, 6\u2014nitrocellulose slurry inlet, 7\u2014NC slurry outlet, 8\u2014blade, and 9\u2014blade holder",
        "texts": [
            " During the process of fine-breaking, the addition of a certain amount of Na2CO3 is required to neutralize the acid. (2) Continuous fine-breaking machine 1. Cone-shaped chopper Cone-shaped chopper is composed of some units, which has a series of advantages, such as large production capacity, relatively stable quality, high efficiency, small site area, and low labor intensity. On the other hand, it has some shortcoming, including complex equipment structure, expensive maintenance, high motor power, and so on. The structure of the cone-shaped chopper is shown in Fig. 10.37. The cone-shaped chopper consists of a rotor and a stator. The rotor is a cone shape, in which the top is installed with blades along the axial. There are several groups of blades with six blades in each group with a knife thickness of 3\u20135 mm. Each blade has a certain gap between blades while the gap between groups is wider. The stator is a sleeve located outside the rotor, fixed to the case. There are equipped with two groups of the A-shaped knife on the stator along the axial. The taper of the vertebral body is 16\u00b0 32\u2032"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure20-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure20-1.png",
        "caption": "Fig. 20. Quick-Emptying Diaphragm Valve",
        "texts": [
            " Temperature effects are a factor to be reckoned with in the operation of both the suction scoop tube and the ejector method of filling, hence further schemes for regulating the coupling without auxiliaries were studied. Gravity Filling and Rim Pumping. One method, illustrated by Fig. 14, involved reversion to the old practice of gravity filling, the scoop tube being dispensed with in this case and provision made for very rapid emptying. The fluid coupling was arranged in a casing, of which the bottom part formed a shallow sump and the upper part a gravity tank immediately above the coupling. Quick-emptying valves were provided, of the kind shown by Fig. 20, which are held shut by the centrifugal head of liquid when the coupling is running. They are 7 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 99 opened simply by cutting off the feed from the gravity tank so that the working circuit quickly empties. The rim of the coupling was embraced at the lower part of its circumference by a stationary chute so that when the emptying valves opened the rush of oil from the coupling was directed upwards by the chute into the gravity tank",
            " This operation was not easy with the pump- controlled coupling because when the runner is stalled a considerable flow of liquid is forced out of the working circuit to the external reservoir and must be pumped back again. With the scoop-controlled coupling, however, this liquid is returned to the working circuit instantly by the scoop tube and the torque can therefore be quickly varied or accurately held. (10) For exceptionally rapid emptying, the leak-off nozzles may be 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 106 PROBLEMS OF FLUID COUPLINGS replaced by diaphragm valves shown by Fig. 20, which were developed\u201d for quickly emptying fluid couplings associated with Voith torque converters. So long as the scoop tube is feeding liquid to the control port a centrifugal head is maintained on top of the thin diaphragm valve, and the large discharge port from the working circuit is closed. Immediately the scoop tube is withdrawn the pressure above the diaphragm leaks away and the valve opens to permit rapid emptying. Traction Uses. The scoop-controlled coupling with quick-emptying valves has scope for cranes and winches ; also for future use as a traction coupling for self-propelled vehicles, since it is glandless, and imposes no idling drag when empty",
            " Sykes had been developed without outside assistance and in his opinion it possessed nothing but disadvantages compared with the standard designs, which were based upon the collaboration and pooled experience of an international group. The disadvantage of the scoop arrangement shown in Fig. 11 was insurmountable for fan drives, since a scoop tube in the position shown could not possibly transfer liquid from the working circuit to effect any useful range of speed control. The diaphragm type of quick-emptying valve shown in Fig. 20 was closed by admitting liquid on the top or radially outermost side of the diaphragm, which was of larger area than the under side exposed to the pressure of liquid in the working circuit. T o open the diaphragm valve the supply of liquid to the top side was shut off, with the result that the small quantity of liquid in the filling passage leaked off centrifugally through small ports and the internal pressure overcame the control pressure. The valve then opened wide, discharging the contents of the working circuit through large ports in the side of the valve housing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000222_0954406219885979-Figure18-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000222_0954406219885979-Figure18-1.png",
        "caption": "Figure 18. Force analysis of the ball at high speed.",
        "texts": [
            " For a special time t \u00bc !Z, the distribution of balls under y\u00bc 3 is symmetrical to that under y\u00bc 3 at t\u00bc 0. Therefore, when the bearing runs, the stiffness coefficient fluctuates between the two curves corresponding to y\u00bc 3 and y\u00bc 3 . Furthermore, the stiffness coefficients of the double-row self-aligning ball bearing with three tilting angles get closer with the increasing rotary speed. The bearing stiffness softening phenomenon can be explained by a force analysis of the ball at high speed, as shown in Figure 18. The outer contact force Qo, the centrifugal force Fc, the inner contact force Qi, and the friction force 2Mg D are in quasi-static equilibrium state. The centrifugal force and the gyroscopic moment of the ball increase with the rotary speed. As a result, the balls move toward to the mid-plane of the bearing, the outer contact angle decreases while the inner contact angle increases. Under these conditions, ball equilibrium is determined by both the applied loads and the level of the centrifugal force and the gyroscopic moment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000723_iconac.2016.7604949-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000723_iconac.2016.7604949-Figure1-1.png",
        "caption": "Fig. 1 Tail-Sitter UAV",
        "texts": [
            " In order to improve the adaptive ability to changes of external environment, Yang designed an adaptive dual fuzzy controller by adding the external state variables in the system inputs [10]. In order to improve the dynamic response performance and steady state accuracy of the system, and improve the adaptive capacity to the changes of flight state, a variable universe fractal dual fuzzy PID controller is designed to carry out the real-time control of the longitudinal attitude of the tail-sitter UAV. The tail-sitter UAV studied in this paper is shown in Figure 1. Flying wing configuration is used and the airfoil of the UAV is symmetric. There are two brushless motors and two elevons are equipped on the fuselage to control the attitude of the UAV. Supposing the vehicle is rigid body and the acceleration of gravity does not change with the variation of UAV\u2019s altitude. Take ground coordinate system O-XYZ as inertial coordinate system. Assuming that the angle rates of roll direction and yaw direction are zero during the transition, and \u03b1\u03b8 = . The dynamic equation of pitch angle motion is given in [11] 21 c 2 [ ] qem m e m y V S c C C C I V\u03b1 \u03c3 \u03c1 \u03b8 \u03b8 \u03b4 \u03b8= + +  (1) Where V is the air velocity, which is the sum of downwash flow velocity of the propellers and the airspeed",
            " This was because the aerodynamic efficiency of elevons increased a lot, and the parameters of PID controller tuned at low airspeed could not meet the requirements of the flight state at this time. As shown in Table 2, The system overshoot of variable universe fractal dual fuzzy PID controller was the smallest, and the adjusting time was the shortest. When the airspeed was large, C would take a small value to decrease the PID parameters which matching the flight state better, so that the adaptive ability of the controller was improved. Based on the prototype of the tail-sitter UAV in Figure 1, the performance of variable universe fractal dual fuzzy PID controller during the transition was verified, and PID controller was used for comparison. During the experiments, the UAV was converted from hover state to horizontal flight status (see Figure 10). the pitch angle during the transition was shown in Figure 11. As seen in Figure 11(a), oscillation occurred during the transition in the system using PID controller. The parameters of PID controller could not meet the requirements of the flight status at this time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001297_0954405419883052-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001297_0954405419883052-Figure5-1.png",
        "caption": "Figure 5. Diagram of light projection modification based on tool geometry.",
        "texts": [
            " For the cutting process, considering that the cutting direction is variable, the ball-nosed end milling is generally used to keep the reference point of the tool (which lies on the tool axis) unchanged. For the printing process, the structure of the printing tool is relatively complex, so the model needs to be simplified accordingly for the calculation convenience. The printing tool can be simplified as a cylinder with a certain radius and the lower printing head can be simplified as a cone connected with the cylinder. In addition, it is necessary to offset the point source by an angle ar when calculating the TAR, as shown in Figure 5. The calculation method of the offset angle ar is as follows ar =arcsin RT LC \u00f010\u00de where ar denotes the offset angle, and LC denotes the distance between the light source and the contact boundary of the obstacle along the tool direction. The proposed sequence planning method of hybrid manufacturing Under the hybrid alternative additive\u2013subtractive manufacturing mode, it is necessary to divide the part into several elements, that is, subcomponents, so as to define a non-interference and support-free forming process plan for complex structural parts"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002222_j.surfcoat.2020.125371-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002222_j.surfcoat.2020.125371-Figure16-1.png",
        "caption": "Fig. 16. Nucleation and growth of TiC, (b) formation of W2C, (c) formation of the W layer, and (d) formation of the irregular layer.",
        "texts": [
            " 15(c) shows, in the gap between two adjacent TiC grains, liquid Ti contacted the W2C layer, and the W2C+Ti\u2192 2W+TiC reaction occurred when the temperature was below 2364 K. In Fig. 15(d), the TiC reaction products filled the gaps between two adjacent TiC, and a compact regular cellular-shaped TiC layer formed around the WCP. The W reaction product was trapped by the W2C/TiC interface, and isolated W was formed in the regular cellular-shaped reaction layer. The formation of a compact TiC layer around the WCP prevented the liquid Ti from making direct contact with the W2C layer and prevented further degradation of it. In Fig. 16(a), the TiC nucleation rate was low due to the low cooling rate conditions in the middle and bottom of the laser molten pool. Thus, the formation of TiC around the WCP was sparse, and the liquid Ti contacted the W2C layer directly. W2C is a metastable phase, and the W2C+Ti\u2192 2W+TiC reaction could take place at low temperatures (1919 K < T < 2364 K). As shown in Fig. 16(c), a continuous W layer formed on the surface of the W2C layer. Chen et al. showed that after W2C was converted to W, the volume of W decreased, and the introduced tensile stress in the continuous W layer then resulted in cracking of it [19]. Moreover, the surrounding liquid Ti infiltrated the W layer along the cracks and reacted with the W2C to form TiC. As shown in Fig. 16(d), mixed reaction products of W2C/TiC/W were formed in an irregular reaction layer encircling WCP. Finally, the difference in the average thickness of the reaction layer around WCP in the laser-induction hybrid melt injected coatings is larger than that of the laser melt injected coatings. Interfacial reaction is a diffusion process that was determined by the dwell time and peak temperature of the laser molten pool. As listed in Table 5, the longest dwell time and highest temperature of the laser molten pool of pre/ post-LIHMI among the samples herein led to the thickest interfacial reaction layer and excessive dissolution of the WCP"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003884_20.717751-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003884_20.717751-Figure1-1.png",
        "caption": "Fig. 1. Skewed model principle. 1 .a. 3D model, 1.b. Sliced model.",
        "texts": [
            " Consequently, the time computation and storage memory do not increase with the movement consideration. With this model, the unknowns are the edge circulations of a and the armature current phases. B. 2 0 Skewed Slots Model In this approach, the skewed slot effects are taken into account by a discretisation in the third dimension. The skewed part of axial length \"z\" is cut into \"n\" disks from an ideal 2D machine by perpendicular planes to the shaft. Two adjacent disks are rotated by an angle of \"ah\" , where a corresponds to the total angle of the skew (see Fig. 1). The winding and bar currents are assumed to be continuous from disk to disk. For each section, the magnetic field equation is solved by 2D finite element in connection with the electric circuits equations of stator and rotor. In these conditions, for each section defined by the rotating angle ai, the magnetic equation is given by : {vgrad U ' , grad ua, ds+ IC oa'. (a, aa, + AV,, ) ds (5) - 5 a'jods=O 4 1 where Avai represents the potential difference per unit of length and aai the nodal values of magnetic vector potential"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002501_j.engfailanal.2020.104977-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002501_j.engfailanal.2020.104977-Figure4-1.png",
        "caption": "Fig. 4. Testing rig. (a) schematic of the testing rig, (b) photo of testing rig.",
        "texts": [
            " The simulation parameters for the specific pneumatic flexible member are as follows: c = 1000 J\u22c5kg\u2212 1\u22c5K\u2212 1, \u03b7 = 0,188%, \u03bb = 0.0158 W\u22c5mK\u2212 1, \u03b1 = 3.5 W\u22c5m\u2212 2\u22c5K\u2212 1, S = 0.05 m2. A series of experiments have been designed in order to verify the mathematical model of heat generation and transfer in the pneumatic pocket as described in Section 3. A testing rig has been designed and manufactured. A schematic representation and photo of the testing rig for the measurement of temperature of the pneumatic flexible member is shown in Fig. 4(a) and Fig. 4(b). Testing rig consists of a frame (1), in which the pneumatic flexible member is mounted (2), oscillating mechanism (3), an electric motor (4) with continuous frequency control between 0 and 50 Hz. Digital multimeters M-3870D METEX with temperature probe ETP\u2013003, and measurement range \u2212 50 \u25e6C to +250 \u25e6C have been used. Temperature probes were set up in three locations where temperature was measured. The following temperatures were measured: \u25aa Air temperature inside the flexible member Tair \u25aa Temperature of the inner surface of the flexible member Tin J. Krajn\u030ca\u0301k et al. Engineering Failure Analysis 119 (2021) 104977 \u25aa Temperature of the outer surface of the flexible member Tout Testing rig was thermally insulated. In Fig. 4(b), the part of thermal insulation of testing rig is visible. During experiments, a stable temperature was maintained in isolated volume around the testing rig. Temperature of the surroundings (which is the same as all the temperatures throughout the system at time t = 0 s) was T0 = 22 \u25e6C. Sampling resolution was 1 min. Temperatures were recorded at times t = 1, 2, 3, \u2026 30 min at constant frequency and amplitude of oscillations and varying air pressure. During experiments, the influence of the environment was minimized by the use of thermal insulation and by maintaining stable temperature conditions around the testing rig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002161_9783527813872-Figure2.72-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002161_9783527813872-Figure2.72-1.png",
        "caption": "Figure 2.72 (a) Perovskite nanocrystal LED. (b) Energy-level diagram. (c) Electroluminescence spectra of red-, orange-, green-, and blue-emitting perovskite nanocrystal LEDs. (d) Images of perovskite nanocrystal LEDs in operation. Source: Li et al. 2016 [362]. Reproduced with permission of John Wiley & Sons.",
        "texts": [
            " Perovskite nanocrystals are considered a good light source for LED applications because of their high color rendering index and narrow emission band. P\u00e9rez-Prieto and coworkers reported MAPbBr3 perovskite nanocrystal-based LEDs for the first time [344]. However, the LED performances were very poor because of low PLQY (c. 20%). Instead of hybrid perovskites, all-inorganic CsPbX3 nanocrystals have been explored with many different interfacial layers. Richard Friend and coworkers developed new cross-linking materials to fabricate efficient perovskite LEDs [362]. They used ITO/CsPbX3 QDs/TFB/MoO3/Ag structure (Figure 2.72). A new TMA vapor phase was used to cross-link the nanocrystal films. Therefore, they were able to coat TFB polymer (poly[(9,9-dioctylfluorenyl-2,7-diyl)-co-(4,4\u2032-(N-(4-sec-butylphenyl)di- 156 2 Synthesis of Solution-Processable Nanoparticles of Inorganic Semiconductors phenylamine)]) on the perovskite nanocrystal layer. They successfully improved luminescence intensity up to 2000 cd/m2 and EQE to 5.7%. Also, Zeng and colleagues reported that a polymeric mixture of perovskite nanocrystals can be used for light conversion materials for white LED [363]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001209_jfm.2019.681-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001209_jfm.2019.681-Figure8-1.png",
        "caption": "FIGURE 8. (Colour online) Spatial distributions of the normalized exceptional maximum stored energy density Eexc in a vibrating isotropic gold nanosphere (b= 40 nm) and in the surrounding glycerol\u2013water mixture for the breathing mode n= 0 using the linear Maxwell model (a,c) and the compressional non-Newtonian model (b,d): (a,b) \u03c7 = 0 (pure water); and (c,d) \u03c7 = 0.8. The colour scale represents the exceptional maximum stored energy density normalized by the energy density at the outer surface of the nanosphere. The radial thickness of the fluid computational domain is limited within 2b.",
        "texts": [
            " The difference of the quality factor of the second class of vibration predicted by these two fluid models is plotted in figure 7 for different glycerol mass fractions. It is found that, for low mass fractions, the compressional relaxation still has a noticeable influence on the breathing mode n= 0 because of its compressional motion behaviour, but its effect on the higher-order modes is less significant compared with the breathing mode and is negligible. For high mass fractions, the vibration characteristics strongly depend upon the compressional relaxation process over the whole mode range. Moreover, for different glycerol\u2013water mixtures, figure 8 depicts the spatial distributions of the normalized exceptional maximum stored energy density Eexc (see SM-IV for more details) in the vibrating gold nanosphere as well as in the surrounding fluid for the breathing mode n = 0 predicted by the linear Maxwell model and the compressional non-Newtonian model. Since the breathing mode has only the radial displacement component (which is independent of \u03b8 and \u03d5), the spatial distribution of the energy density only varies along the r-direction and does not change along the \u03b8 -direction, as shown in figure 8. One can observe that the compressional relaxation effect on the energy density distribution of the breathing mode is more significant for a high glycerol mass fraction (\u03c7 = 0.8) than a low mass fraction (\u03c7 = 0). In particular, since the breathing mode is a spherically symmetric compressional motion, ht tp s: // do i.o rg /1 0. 10 17 /jf m .2 01 9. 68 1 D ow nl oa de d fr om h tt ps :// w w w .c am br id ge .o rg /c or e. A cc es s pa id b y th e U CS F Li br ar y, o n 05 O ct 2 01 9 at 1 3: 11 :4 6, s ub je ct to th e Ca m br id ge C or e te rm s of u se , a va ila bl e at h tt ps :// w w w "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000876_0954407018824943-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000876_0954407018824943-Figure1-1.png",
        "caption": "Figure 1. FEAD with its poly-V serpentine belt drive. FEAD: front engine accessory drive.",
        "texts": [
            " The poly-V belt is a composite made of elastomeric material layers, generally EPDM (ethylene propylene diene monomer) rubber, and a cord layer (poly ethylene terephthalate (PET), fibre). Flat on its top surface (backside), several ribs compose the opposite side of the poly-V belt to increase the contact surface with the pulleys. Idler-pulleys are often used in this type of transmission to increase the wrap angle around the pulleys. A tensioner pulley is used to adjust the belt tension according to the various operating conditions of the engine (Figure 1). Nowadays, reducing the power losses of engines has become a design matter. Most of the works related to power losses in belt transmission were initiated by Childs et al.1,2 and Gerbert.3,4 However, only flat and V belts were considered. Almeida and Greenberg5 and Chen et al.6 studied the global efficiency of belt transmission and therefore the global power loss. In the case of poly-V belts, experimental studies were carried out on the analysis of the belt\u2013pulleys slip7 and on the pulley\u2013belt friction coefficient identification",
            " To go further, new models to predict the power losses due to belt tension fluctuations and tensioner-hysteresis are given. Finally, numerical results are generated for several test bench cases in order to evaluate the developed power loss models: the belt-hysteresis, the pulley\u2013belt slip and the bearing power losses. The numerical results are compared with experimental ones. The FEAD representative Power Loss (PLFEAD) model is defined by equation (1) PLFEAD =PLhys +PLbear +PLslip +PLvib \u00f01\u00de The power loss model PLFEAD represented by equation (1) suggests that in a poly-V belt transmission as in Figure 1, energy is dissipated by 1. Hysteresis PLhys, 2. Friction inside the bearings PLbear, 3. Slip between the poly-V belt and the pulleys PLslip, 4. Vibrations in the belt-spans PLvib. In equation (1), one part of the energy is dissipated inside the belt due to the hysteretic behaviour of its constitutive elastomer PLhys belt, and another part, dissipated outside the belt due to the tensioner hysteretic behaviour PLhys tens. In equation (4), belt-hysteresis power losses are induced by tension fluctuations in the belt-free and wrapped spans of the FEAD",
            " Rotational motions are induced by the engine acyclism as given, for example, by equation (35) which corresponds to a truck engine case. In this study, the belt tensions and the tensioner angle fluctuations are considered in this paper as periodic and correspond to a dynamic stationary state vCS t\u00f0 \u00de= 600+16cos 2p30t\u00f0 \u00de\u00bd 2p 60 \u00f035\u00de Equation (35) represents the angular speed fluctuations with a single cosine; however, a multi-harmonic engine acyclism can longitudinally excite the poly-V belt when the truck engine (Figure 1) is not idling. It means that vCS . 600r=min in equation (35). The tension fluctuations resulting from the excitation in equation (35) are determined through numerical integration of equations of motion17,18 or using modal superposition.16 The non-linearity due to potential periodical stick slip motion of the tensioner22 is not considered in the present analysis. The objective is to determine the periodic rotation angle of the tensionerarm and the periodic fluctuations of the belt-span tensions",
            " Changing from situation S1 to S2, the objectives are (1) to validate globally the power loss models in equation (1) and (2) to highlight the effectiveness of the PLslip modelling, since except PLslip, the other models in equation (1) have already been either analysed, for example, PLhys, 10,11 or assumed, by default, to be validated, for example, PLbear from SKF.13 The belt used here has the same characteristics as that in Silva et al. except its length (Table 2). Moreover, it is installed with a setting-tension T0 =204N which is far below a common setting-tension representative of Truck FEAD applications (T0 =600N) (Figure 1). However, as previously mentioned, a lower settingtension permits highlighting new results (PLslip) not verified yet in the previous analyses.10,11 The FEAD power losses represented by equation (1) were both measured and simulated, with the model presented, see Table 3. Since the test bench in Figure 14 is equipped with a speed and torque measuring system on each shaft, the power supplied by the CS and the power consumed by the accessory drives are known. The difference between the power supplied and consumed is the power lost/measured in the FEAD, PLM in Table 3",
            " PLFEAD is constant and independent of the transmitted power; (B) where T0 starts becoming insufficient and the PLFEAD increases linearly with the transmitted torque/power; (Gross Slip) where T0 is not anymore enough to transmit the desired power and PLFEAD increases exponentially. Moreover, note that gross slip only occurs experimentally and it is neither considered by the model in equation (1) nor partially by Childs in his PLslip model. It is because to predict analytically the power loss PLFEAD in the gross slip zone is hard (due to complex tribological effects) and unnecessary since it never corresponds to the desired operation point of an engine as in Figure 1 working properly. Finally, the relative error (PLM\u2013PLS) / PLM increases up to a maximum of 5.9% within the engine working conditions (zones A and B in Figure 15); for both measured and simulated power losses, this is a reasonable error which permits validating the power loss models PLbelt hys, PLbear and PLslip in equation (1). The slight discrepancy between measured and simulated power losses in Figure 15 and Table 3 can be explained as possibly: (1) not all existing power loss phenomena have been taken into account; (2) the coefficients used to compute the frictional moments of the bearings might be under/miss estimated; (3) since the temperature in each bearing rolling contact surface is hard to be measured and it directly affects the estimation of bearing losses, the global bearing temperature of 20 C was measured and used in the bearing power loss model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002349_j.jmapro.2020.05.052-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002349_j.jmapro.2020.05.052-Figure1-1.png",
        "caption": "Fig. 1. Slicing process for additive manufacturing using laser-DED technique.",
        "texts": [
            " A predetermined amount of feedstock powdered material, either powder or wire, is feed to a heat source (usually a laser) to form a melt pool which gets deposited on a metallic base or substrate. The molten material is rapidly solidified as the substrate (or deposition head) moves away from that point. The motion of the substrate is controlled using a CNC controller by a predetermined toolpath. This toolpath consists of the build part being deconstructed into numerous layers and hatches as seen in Fig. 1. Hatches can be defined as the tool-paths in the XY direction or in the plane perpendicular to the build direction. This plane is also referred to as build plane. The motion controller moves the substrate in a rasterscanning fashion to deposit individual hatches. The hatches collectively become a layer. A schematic diagram shows the hatch and contour scans in Fig. 1. The deposition head then moves up by a distance equal to the layer thickness to repeat the process to deposit another layer of material via raster-scanned hatches. The process is repeated until the desired build geometry is achieved. A 3D geometry defined by the CAD model is thus manufactured layer by layer. A schematic of a typical laser-DED setup is shown in Fig. 2. The deposition head is enclosed in an argon filled glovebox. The feedstock powder is loaded in an argon pressurized powder feeder outside of the glovebox as seen in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001019_rpj-07-2018-0182-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001019_rpj-07-2018-0182-Figure14-1.png",
        "caption": "Figure 14 Sectioned view of redesigned die",
        "texts": [
            " This result shows that the die is safer when it is designed for hydroforming convolution width \u201cW\u201d 6.5 mm and the applied forming pressure is 9 MPa. From the stress distribution plot, it can also be observed that the stress is higher at the bottom of convolution forming area and lower at other regions. The flange region other than the forming area and vertical guide wall below the convolution show very low stress as shown in Figure 13. These two regions were selected for manual material reduction. Die is redesigned using Solidworks software as shown in Figure 14. The height of the vertical guide wall is reduced as its functions only as a guide for locating the die while fixing. Low stress regions in the flange were redesigned with unidirectional square pores which have nearly optimized compression strength, flexural strength, modulus and build time compared with solid part (Iyibilgin et al., 2014). The square pore of 2.24 mm edge length was selected based on the previous study (Sathishkumar et al., 2016) and minimum wall thickness to create unidirectional pores structures with better homogeneity in the FDMmachine"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002329_042033-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002329_042033-Figure6-1.png",
        "caption": "Figure 6. Rollover of various configurations of the machine-tractor unit.",
        "texts": [
            " This can be explained by such factors as: different tire pressure, error in weighing, spatial distribution of unaccounted masses, track mismatch, difference in stiffness of real and virtual wheels, and settings for the interaction of the contact pair \u201cplatform \u2013 wheels\u201d. Using a simulation model of a tractor, it is possible to study the critical angles of lateral stability of various layout options of a machine-tractor unit with modular attachments [16]. We give an investigation of three typical configurations (figure 6). An analysis of the simulation data of the rollover process of a tractor equipped with a single-beam rear mounted implement (0 + 1 scheme) shows that the contact force of the front wheel begins to decrease immediately. The contact force of the rear wheel after a slight increase also decreases with an increase in the angle of inclination. The first was the separation of the rear wheel at an angle of inclination of 36\u00b0. The separation of the front wheel occurred at an angle of 41 \u00b0. When rollover a tractor equipped with a double-beam rear mounted implement (scheme 0 + 2), the contact force of the front wheels begins to decrease immediately"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002430_tie.2020.3014570-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002430_tie.2020.3014570-Figure15-1.png",
        "caption": "Fig. 15 Magnetic flux leakage of the slot opening and parameters of the curved tooth for optimization.",
        "texts": [
            " The local magnetic saturation problem of the stator tooth caused by the armature reaction is the main reason for the asymmetry of torque waveform about \u03b8=0 [7]. Fig. 14 shows the magnetic flux leakage of the slot opening and parameters of the curved tooth for optimization. It can be observed from Fig. 14 that the reason mainly comes from an excessive magnetic leakage of slot opening. Increasing the slots leakage reluctances can reduce the saturation problem at tooth-tip segment effectively. Therefore, different from the original straight tooth shape, a curved nonwounded tooth shape of the stator is carried out and shown in Fig. 15. The magnetic leakage of slot opening is suppressed by increased the path length of slot leakage. In order to clearly Fig. 11 Comparison of torque versus current characteristic between the straight and curved tooth sh pe. Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on September 21,2020 at 08:45:24 UTC from IEEE Xplore. Restrictions apply. 0278-0046 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure4-1.png",
        "caption": "Fig. 4. Arbitrary object with a center of mass located at C.",
        "texts": [
            " This section describes our proposed method to estimate all six inertia tensor components and the position of the centre of mass of a rigid-body object. The method can be used with the measurements collected using the torsion platform described in Section 4, or another equivalent platform. The method requires a set of estimates of the mass moment of inertia of the object about a vertical axis (IRzz, defined in this section). In the case of a torsion platform, these are obtained by measuring the period of oscillation of the base platform with the object mounted in a series of specified configurations. The proposed method considers an object (Fig. 4) with its CM located at rC with respect to frame O. Frame O has the axis zO aligned with the axis of rotation of the system (Fig. 8). Frame B is a body fixed frame located at an arbitrary location rB with respect to frame O and is initially parallel to it. Frame C is a body fixed frame with origin at the CM and parallel to frame B. The object has an unknown inertia tensor I with respect to frame C: I \u00bc Ixx Ixy Ixz Ixy Iyy Iyz Ixz Iyz Izz 0 B@ 1 CA \u00f010\u00de Vector rB can be measured directly: rB \u00bc xB; yB; zB\u00f0 \u00deT \u00f011\u00de while the vectors rC=B and rC are unknown: rC=B \u00bc xC=B; yC=B; zC=B T \u00f012\u00de rC \u00bc rB \u00fe rC=B \u00bc xC ; yC ; zC\u00f0 \u00deT \u00f013\u00de The inertia of the object prior to any rotations can be determined relative to frame O from the parallel axis theorem: IO \u00bc I \u00fem rTCrCI3 rCrTC \u00f014\u00de where I3 is a 3 by 3 identity matrix, and m is the mass of the object"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003645_s0166-1280(98)00321-2-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003645_s0166-1280(98)00321-2-Figure5-1.png",
        "caption": "Fig. 5. The decomposition where scission of the Ca\u2013O bond is rate-determining.",
        "texts": [
            " Models were assembled by one of the alkoxy groups generated from the primary (methanol, ethanol, n-propanol, n-butanol), the secondary (s-butanol, 2-propanol) or the tertiary alcohol (t-butanol), some Oss and one Si atom and were used for calculations of the activation energy for the decomposition of the alkoxy groups. For the decomposition of the alkoxy groups where the extraction of Hb by the oxygen atom of the group, as illustrated in Fig. 4 (hereafter the case of 2-propanol is shown as the example in Figs 4\u20136), is ratedetermining, the heat of formation was calculated in the same way. In order to examine the decomposition of the alkoxy group where scission of the Ca\u2013O bond (Fig. 5) or scission of the Ca\u2013O bond of the protonated alkoxy group (Fig. 6) is rate-determining, the heats of formation were calculated in the same way. In the latter case, the formal charge of the model was set to a unit positive charge because of the attached proton. Calculated results for the case where the dehydration occurs by scission of Ca\u2013O bond of the isopropoxy group on SiO2 are shown in Fig. 7. The heat of formation of the model was calculated when separating the Ca from the oxygen in the group at 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000604_j.cja.2016.06.014-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000604_j.cja.2016.06.014-Figure1-1.png",
        "caption": "Fig. 1 2D wing model with control surface.",
        "texts": [
            " Numerical simulations are given in Section 4. Section 5 briefs the conclusions of the research. In this section, we briefly recall the mathematical model for the flutter of a reentry vehicle with actuators fault-free. Based on this nominal flutter system, the state equation with actuator faults and saturation, parameter uncertainties and external disturbances are established. 2.1. Wing flutter model under actuators fault-free In this section, flutter problem for a 2D wing including cubic hard spring nonlinearity is analyzed. As shown in Fig. 1, a two degree-of-freedom (2-DOF) wing system model is considered herein. The plunge deflection is denoted by h, positive in the downward direction; h is the pitch angle about the elastic axis, positive nose up; V denotes the air speed; the chord length is c; Q, p and C are the aerodynamic center, elastic axis and center of mass, respectively; the distance from the leading edge to the elastic axis is xp, and that from the leading edge to the mass center is xC; dLEout and dLEin (or dREout and dREin) are the control surface angles. From Fig. 1, the velocity of mass center of wing can be expressed as _z \u00bc _h\u00fe \u00f0xC xp\u00de _h \u00f01\u00de The kinetic energy, potential energy and dissipation of the system can be given by T \u00bc 1 2 mW _z2 \u00fe 1 2 me _h2 \u00fe 1 2 IC _h2 U \u00bc 1 2 Khh 2 \u00fe 1 2 Khh 2 f \u00bc 1 2 Ch _h2 \u00fe 1 2 Ch _h2 8>< >: \u00f02\u00de where IC, mW, me, Kh, Kh, Ch and Ch are the moment of inertia about center of mass, wing mass, wing extra-mass, stiffness coefficient in plunge, torsion stiffness coefficient, damping coefficient in plunge and torsion damping coefficient, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002731_j.mechmachtheory.2021.104285-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002731_j.mechmachtheory.2021.104285-Figure5-1.png",
        "caption": "Fig. 5. Geometrical parameters for the gear rim deformation.",
        "texts": [
            " The deformation of the gear body also has a great influence on the mesh stiffness. The deformation of the gear body can be expressed as Eq. (64) [ 28 , 30 ]. \u03b4 f = F cos 2 \u03b1m W E ( L \u2217 ( u f S f )2 + M \u2217 ( u f S f ) + P \u2217 ( 1 + Q \u2217tan 2 \u03b1m )) (64) The coefficients L \u2217, M \u2217, P \u2217, Q \u2217 can be approached by polynomial functions X i \u2217(h f i , \u03b8 f ) = A i \u03b82 f + B i h 2 f i + C i h f i \u03b8 f + D i \u03b8 f + E i h f i + F i (65) the values of A i ,B i ,C i ,D i ,E i , and F i are given by reference [30] . The parameters u f ,s f ,h fi = r f / r int and \u03b8 f are shown in Fig. 5 . Then the equivalent stiffness on the meshing line caused by the deformation of the gear body is 1 K f = \u03b4 f F (66) Based on the above, the mesh stiffness K g of helical gear can be obtained by Eq. (67) . K g = N \u2211 j = 1 ( K e ) j (67) The mesh stiffness ( K e ) j of the j th thin gear slice can be expressed as [31] \u23a7 \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23a9 ( K e ) j = 1 / ( 1 ( K t1 ) j + 1 ( K t2 ) j + 1 ( K h ) j ) ( K t1 ) j = 1 / ( 1 ( K b1 ) j + 1 ( K s 1 ) j + 1 ( K a 1 ) j + 1 ( K f 1 ) j ) ( K t2 ) j = 1 / ( 1 ( K b2 ) j + 1 ( K s 2 ) j + 1 ( K a 2 ) j + 1 ( K f 2 ) j ) (68) where \"1 \u2032\u2032 and \"2 \u2032\u2032 denote the driving gear and driven gear respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000204_j.autcon.2019.102996-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000204_j.autcon.2019.102996-Figure13-1.png",
        "caption": "Fig. 13. Twist angle measurement: (a) numbered components in global view; (b) numbered elements in local view, twist angle of 5.5\u00b0.",
        "texts": [
            " For instance, a vertex may have three connecting lines, while the next vertex of the list may have six. This data structure \u2013 lines grouped by vertex \u2013 shows how many elements connect to one node. Manipulation of the data structure allows efficient measurements of angles and distances between structural components. The twist angle is the angle between one hub longitudinal axis and the other hub axis connected to the same element. Hubs axes are positioned according to the mesh normal vectors, as described in Fig. 13. There is one twist angle for every element. The data is organized in vertices grouped by line for this computation. Each line has node 0 and node 1. The normal vectors are generated through a mesh offset component, as observed in Fig. 13a, in the zoomed circle. Mesh normal vectors (or hub axis vectors) are named hi, i ranging from 0 to the number of nodes in the mesh. Elements are positioned according to the mesh lines. Element direction vectors are named bk, k ranging from 0 to the number of elements in the mesh. Twist angles are computed on the plane normal to bk vectors. A plane perpendicular to b0 is created for the example on Fig. 13b. The twist angle is calculated through a center point, node 0, and two projected points on the created plane. This angle is 5.5\u00b0 in the example of Fig. 13b. The interconnecting parts connect to elements through a screw for the proof of concept model (Fig. 4a). For ease of fabrication, we decided that elements have aligned holes. It would be unpractical to measure the twist angle to drill the holes in the correct angles for each element, which already have different lengths. We included this parameter in the geometry of the interconnecting parts. The adjacency angle is defined as the angle between the elements projected on the plane normal to each hub axis vector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000634_s12555-015-0089-9-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000634_s12555-015-0089-9-Figure1-1.png",
        "caption": "Fig. 1. Description of the related frames and vectors.",
        "texts": [
            " Moreover, \u2225a\u2225 \u2264 \u2225a\u22251, where \u2225a\u2225 and \u2225a\u22251 denote vector 2-norm and 1- norm, respectively. A > 0 denotes A is a positive definite matrix, and \u2225A\u2225 is the induced 2-norm of matrix A. In and On are n\u00d7n unit and zero matrices, respectively. sgn(c)\u225c [sgn(c1), \u00b7 \u00b7 \u00b7 ,sgn(cn)] T for c= [c1, \u00b7 \u00b7 \u00b7 ,cn] T \u2208 Rn, where sgn(ci) =  \u22121, ci < 0 0, ci = 0 1, ci > 0 , i = 1, \u00b7 \u00b7 \u00b7 ,n. 2.1. Dynamics of chaser and target The control problem that a chaser subject to uncertain inertia and unknown disturbance tracks a tumbling non-cooperative space target is considered in this work. Fig. 1 depicts the related frames and vectors, where Fi \u225c {Oxiyizi} represents an Earth-centered inertia frame, Fc \u225c {Cxyz} and Ft \u225c {Txtytzt} are the chaser and target body-fixed frames; O, C, and T are the centers of mass of the Earth, chaser, and target, respectively; P is the chaser\u2019s desired proximity position along the direction of target\u2019s docking port; {r,re} and {rt ,pt ,rpt} are the related position vectors represented in frame Fc and frame Ft , respectively. The aim of this paper is to design controller such that the center of mass C tracks point P and frame Fc tracks frame Ft ",
            " Remark 1: For the target\u2019s dynamics (7) and (8), denote target\u2019s kinetic energy E(t) = 1 2 (mtv T t vt + \u03c9T t Jt\u03c9t) \u2265 0 and from its time derivative E\u0307(t) = \u2212mtv T t S(\u03c9t)vt \u2212 \u03c9T t S(\u03c9t)Jt\u03c9t \u2261 0, we know E(t) \u2261 E(0) \u225c 1 2 [mtv T t (0)vt(0)+\u03c9T t (0)Jt\u03c9t(0)] < \u221e. Thus, the target\u2019s linear velocity vt and angular velocity \u03c9t are always bounded. 2.2. Modeling of relative motion dynamics Define the relative attitude in terms of MRPs as [24] \u03c3e = \u03c3t(\u03c3 T\u03c3\u22121)+\u03c3(1\u2212\u03c3T t \u03c3t)\u22122S(\u03c3t)\u03c3 1+\u03c3T t \u03c3t\u03c3T\u03c3+2\u03c3T t \u03c3 (9) and denote the rotation matrix from Ft to Fc as R= I3\u2212 4(1\u2212\u03c3T e \u03c3e) (1+\u03c3T e \u03c3e)2 S(\u03c3e)+ 8 (1+\u03c3T e \u03c3e)2 S2(\u03c3e). (10) In the light of Fig. 1, the position and velocity of point P with respect to O denoted in frame Ft are rpt = rt +pt , vpt = vt +S(\u03c9t)pt , (11) where the pt \u2208 R3 is a constant vector denoted in frame Ft . The relative position, linear and angular velocities are also denoted in frame Fc by re = r\u2212Rrpt , ve = v\u2212Rvpt , \u03c9e =\u03c9\u2212R\u03c9t (12) Substituting (12) into (1), (2), (3), (4) and using the identities R\u0307 = \u2212S(\u03c9e)R, r\u0307pt = vpt \u2212 S(\u03c9t)rpt and R\u22121 = RT yield the relative motion models represented in frame Fc as [25] r\u0307e = ve \u2212S(\u03c9)re, (13) \u03c3\u0307e = G(\u03c3e)\u03c9e, (14) mv\u0307e =\u2212m[S(\u03c9)v+Rv\u0307pt \u2212S(\u03c9e)(v\u2212ve)]+f +d f , (15) J\u03c9\u0307e =\u2212S(\u03c9)J\u03c9\u2212 J[R\u03c9\u0307t +S(\u03c9)\u03c9e]+\u03c4 +d\u03c4 , (16) where G(\u03c3e) = 1 4 [(1\u2212\u03c3T e \u03c3e)I3 +2S(\u03c3e)+2\u03c3e\u03c3 T e ] is invertible, Rv\u0307pt can be derived from (11), (7), (12) and RS(a) = S(Ra)R for any a \u2208 R3 as Rv\u0307pt =R[v\u0307t +S(\u03c9\u0307t)pt ] =\u2212RS(\u03c9t)vt \u2212RS(pt)\u03c9\u0307t =\u2212S(R\u03c9t)[Rvpt \u2212RS(\u03c9t)pt ]\u2212RS(pt)\u03c9\u0307t =\u2212S(\u03c9\u2212\u03c9e)[v\u2212ve \u2212S(\u03c9\u2212\u03c9e)Rpt ] \u2212RS(pt)\u03c9\u0307t (17) and \u03c9\u0307t can be derived from (8) and (12) as \u03c9\u0307t =\u2212 J\u22121 t S(\u03c9t)Jt\u03c9t =\u2212 J\u22121 t S(RT(\u03c9\u2212\u03c9e))JtRT(\u03c9\u2212\u03c9e)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001900_s0263574720000806-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001900_s0263574720000806-Figure1-1.png",
        "caption": "Fig. 1. Robot with a limited field-of-view in an unknown environment with fixed and moving objects. The black region is unobserved and may contain unexpected objects.",
        "texts": [
            " E-mail: s_bouraine@yahoo.fr https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574720000806 Downloaded from https://www.cambridge.org/core. Cornell University Library, on 12 Sep 2020 at 17:55:55, subject to the Cambridge Core terms of use, available at limitations (presence of unseen objects in the limits of the robot\u2019s field-of-view and occluded regions) and the second one is related to the dynamics of the environment (presence of moving objects with unknown future behaviour) (see Fig. 1). The purpose of this paper is precisely to propose a motion planning method that handles such challenging constraints using an appropriate state\u00d7 time model of the future, provides strict guarantees of motion safety with formal proofs and solves the trajectory optimization problem by proposing a technique that formally integrates safety to compute an optimal trajectory to the goal. Given that motion safety has to do with staying away from states where a collision occurs (now or eventually), the first position taken in this work is to address the motion safety issue. In an ideal case, motion safety is the guarantee that no collision between the robot and its surroundings will ever occur whatever happens. Theoretically, this form of safety (called absolute motion safety) consists in finding a collision-free trajectory of infinite duration, which requires knowledge of the future up to infinity (a priori known objects behaviour). In situations such as Fig. 1, this form of safety is impossible to guarantee.1 In a previous work,2 this problem was solved by considering a weaker level of safety (better guarantee less than nothing) by guaranteeing that the robot will be at rest before an inevitable collision occurs whatever happens. A passive motion safety is therefore guaranteed where the robot takes its own responsibility with respect to the collision problem by braking and stopping before collision occurs. This choice was provably argued in ref. [2] the proposed solution is based on the braking inevitable collision state concept [braking inevitable collision states (ICS)] (for more details, see refs",
            " Passive motion safety is obtained by avoiding braking ICS at all times. Unlike existing trajectory optimization approaches, the proposed approach PASSPMP-PSO has the ability to generate near-optimal trajectories when staying away from braking ICS whatever happens, that is, near-optimal passively safe trajectories. The main contributions and novelties of this paper are the following: (1) We propose a new motion planner dubbed PASSPMP-PSO that integrates both safety guarantee and trajectory optimization in challenging scenarios as in Fig. 1. Among the few works rigorously addressing the safety issue (e.g., refs. [9,14]), none of them deals with trajectory optimization even if the two issues are strongly related (both are necessary for a good performance of a mission). (2) PASSPMP-PSO computes a passively safe near-optimal partial trajectory, where motion safety and trajectory optimization are addressed in a formal way. PASSPMP-PSO reasons about the future evolution of the environment over a limited lookahead, which is formally defined so as to ensure passive safety guarantee",
            " At each cycle, a near-optimal and passively safe partial trajectory is generated. It is guaranteed that at each cycle, a p-safe trajectory is always found. PASSPMP-PSO algorithm has been implemented in simulation and in real-world experiments, to illustrate its performances, respectively, in challenging scenarios and real situations. Furthermore, it has been evaluated wrt two other motion planning methods. 5.1.1. Test conditions. To validate PASSPMP-PSO and demonstrate its performances, it has been implemented and tested in simulation on scenarios similar to that of Fig. 1. The simulation environment is a 2D workspace of 180\u00d7 180 m, featuring fixed and moving objects with arbitrary trajectories2 and an upper bounded velocity vbmax = 20 m/s. The used robotic system A is a car-like vehicle, governed by the following dynamics:\u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x\u0307 y\u0307 \u03b8\u0307 v\u0307 \u03be\u0307 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6= \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 v cos \u03b8 v sin \u03b8 v tan \u03be/L 0 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6+ \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 0 0 1 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 u\u03b1 + \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 u\u03be (22) where A\u2019s state is a 5-tuple s = (x, y, \u03b8, v, \u03be), (x, y) the Cartesian coordinates, \u03b8 the orientation, v the linear velocity and \u03be the steering angle",
            "49 Among planning approaches, this approach is the most suitable for unknown environments where the robot only uses its on-board sensors for perception. Furthermore, it is very adapted for real-time applications and it is highly applied in road networks and rough terrains. This method is used as a local planner, thus gaining in reactivity. However, the planned trajectories can be biased by a global planner so as to guide the robot towards the goal. Therefore, this method can be easily adapted to such constraints of Fig. 1. These constraints are the most challenging with regard to the safety guarantee criterion, which is the main purpose of PASSPMP-PSO. ISS is a straightforward approach intended to generate feasible motion planning search space, where all motions sampled in the input space are inherently executable. Each input is simulated using the predictive motion model (dynamics of the robotic system) to determine the shape of the resulting trajectory. The generated trajectories can be evaluated according to a given cost function and can be tested for collision-free"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003868_9.486641-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003868_9.486641-Figure4-1.png",
        "caption": "Fig. 4. A state partition and a continuous state trajectory of the hybrid system.",
        "texts": [
            " If d ( k ) = 0, then the robot performs its task as before, but if d ( k ) # 0, then the next state may not be as before. The state trajectory of the disturbed system is shown in Fig. 3(b). As is shown, it takes more steps to reach XO, hut the robot performs its task properly, despite a constantly changing disturbance . B. A Double-Integrator Example Consider the CSS in Fig. 1 as a double integrator given by = XZ, iz = U , where U can take values from the set U = { - l , O , +I}. The state partition prescribed by an interface is shown in Fig. 4. The corresponding nondeterministic system is X ( k + 1) E F ( X ( k ) . U ( k ) ) , where X ( k ) = [Xl(k) Xz(k)]\u2019 and U ( k ) E M. As an example, observe that F ( [ 2 2]\u2019, -1) = {[2 l]\u2019}, and F ( [ l 11\u2019. -1) = {[I 01\u2018. [2 l]\u2018, [2 O]\u2018}. As is shown, for X ( k ) = [2 2]\u2018, the next state is deterministic (that is F has a single element), and for X ( k ) = [l 11\u2019 with the input -1, the next state is not deterministic. Consider a controller of the form U ( k ) = -sign(Xl(k) + X z ( k ) ) . (31) For z(0) = [1.5 1.5]\u2019, the continuous state trajectory of the system is shown in Fig. 4. Although no high-frequency switchings occur in the DSS, the sliding mode is in effect by definition along the lines shown in Fig. 4, since many other trajectories result in reaching the same manifold. We note that the discrete-state trajectory reaches X = [0 01\u2019. Since we use the partitioning above, for awhile the trajectory leaves the state [0 O]\u2019, but the control makes the trajectory reach it again. A finer partitioning of the state space can result in more precise trajectory regulation. VII. CONCLUSIONS In this paper, dynamic systems with discrete and hybrid state space are considered. In some cases classical stability is not adequate to directly apply to DSS\u2019s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure26-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure26-1.png",
        "caption": "Fig. 26. (a) Simulated casing wear under 500RPM and 200 bar; (b) casing wear detected on the interior wall of the pump housing.",
        "texts": [
            " Also, it is necessary to note: although body dynamics may be less important in predicting volumetric efficiencies when compared to a good prediction for the mean gear position, it is crucial in predicting the balancing, vibrations, and wear, as discussed in the earlier sections in this paper. The main strength of the model discussed in this paper is to model the micro-movements and balancing of the major solid components of a CCHGP-type pump, however, the measurements of the forces, moments, and micro-motions are not possible at this time, which can be expected in the future work. Nevertheless, this simulation model is successful in replicating the casing wear pattern found in the tested pump, as shown in Fig. 26, where the predicted casing wear is produced under a 500RPM, 200Bar low-speed and highpressure operating condition. As extensively discussed in Section 6.1, the complicated 3D loadings on the rotors given by its complex geometry result in tilting of the rotors. Unless properly balanced, casing wear in 3D fashion could show up as a result of the complex 3D gear dynamics. The photo in Fig. 26b validates this consideration, in which a strip of wear on the driver gear side of the internal casing can be clearly seen. This casing wear pattern is accurately replicated by the simulation model (Fig. 26a), which prominently showcases the potential of the current model in predictive modeling and design optimizations of balancing mechanism. Continuous-Contact Helical Gear Pump (CCHGP) is a novel hydraulic gear pump design, featuring low-fluctuation/no-fluctuation fluid delivery, and extraordinary low noise emission level. Such merits are given by its special rotor geometry, which nevertheless also entails complex 3D loading patterns and resulting 3D gear dynamics. This paper, together with its preceding work [20], presents a multi-domain dynamic simulation model for the CCHGP type of pumps, which is validated by the experimental measurements and observations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002085_tec.2020.3044000-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002085_tec.2020.3044000-Figure14-1.png",
        "caption": "Fig. 14. Prototype 39-slot and 12-pole motor with UBW structure",
        "texts": [
            "org/publications_standards/publications/rights/index.html for more information. 9 can be seen that the UMP force occurring on the rotor is zero for the conventional 36-slot motor since the winding distribution is symmetric. A prototype motor with unbalanced winding and 39-slots and 12-poles has been manufactured to verify the motor performance results experimentally. The specifications of the prototype machine are given in Table VII [19]. The stator and rotor laminations and mounted un-skewed rotor structure are illustrated in Fig. 14. In addition, 2-dimensional (2D) finite element analysis of the same motor design was performed to obtain both no-load and on-load motor performance results. Then, FEA and experimental results are compared. The comparison of line-to-line back-EMF voltage waveforms at 1000 rpm between FEA and test results is shown in Fig. 15(a). It is seen that the FEA results are in good agreement with the test results. There is only 0.9% discrepancy between the FEA and the experimental results. In addition, line-to-line backEMF voltage waveforms for all phases obtained from test results at 1000 rpm rotor speed are illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure11-1.png",
        "caption": "Fig. 11. Plate and other rotating components.",
        "texts": [
            " The spring stores energy during angular displacement and provides the restoring torque to oscillate the plate assembly. The spring was selected such that the period of oscillation of the plate assembly is at least 1 s. The test object is supported on the rotating plate by customized attachment blocks. The rotation plate was fabricated from 6061-T6 aluminum by water-jet cutting, with a dimensional tolerance of 0.05 mm, and has the dimensions specified in Fig. 10. The components of the rotating assembly are shown in Fig. 11. The rotation plate has a pattern of machined holes that allow blocks to attach in different positions and orientations with respect to the rotation axis. The attachment blocks can be 3D printed to reduce costs of manufacturing, or can be machined and positioned with high tolerance positioning pins to reduce the uncertainty of the position and orientation of the object. A two-degree-of-freedom joint mechanism can also be used instead of several attachment blocks. In that case, the joints may need to be instrumented or have mechanical stops to set the orientation of the object, or the orientation can be measured directly",
            " It is desired to find a region where the random experimental errors have a normal distribution, and the data is free from outliers. The measurements should also be tested to see whether their errors meet the i.i.d. assumption used by linear regression. A set of experiments was conducted to validate the torsional platform apparatus described in Section 4.1. These tests consisted of estimating IOzz of the rotating assembly in six attachment configurations and comparing the estimates to the inertias obtained from a SolidWorks model of this assembly in the same configurations (Fig. 11). The SolidWorks model is in close agreement with the physical assembly; for instance the mass of the rotation plate was reported to be 0.87041 kg and it was measured to be 0.870 kg on a calibrated balance. The components can be expected to have constant densities, therefore, the inertia tensors are expected to be in close agreement. Six different configurations were selected and ten datasets were collected for each configuration following the procedure in Section 4.6. The system was configured by connecting the shaft to the rotating plate in six different locations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000491_s10846-016-0360-1-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000491_s10846-016-0360-1-Figure15-1.png",
        "caption": "Fig. 15 3D positioning system scheme based on OptiTrack device",
        "texts": [
            " For instance, the altitude evolves in a range from 20 cm to 40 cm, it means about 75 % larger than altitude experi- enced in previous experiments when optimal control is used and a similar behavior has the velocity z evolving from \u221210 cm/seg to 10 cm/seg. In order to complement this work, a discrete optimal control strategy without exact linearization is tested and also compared with the optimal control law with exact linearization, both are applied in a path tracking task, and in this case an OptiTrack system (see Fig. 15) is used to measure the 3D position of the UAV. The obtained real-time experiment results are plotted in Figs. 16, 17 and 18. Firstly, the altitude behavior is shown in Fig. 16 when both controllers are implemented. One can realize the experienced behavior is similar for this couple of controllers. However, the optimal control with exact linearization has a lower energy consumption rate than that related to the controller without exact linearization, it can be corroborated by the energy consumption analysis and the results are presented in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000844_1464419318819332-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000844_1464419318819332-Figure2-1.png",
        "caption": "Figure 2. RD interaction with OR.",
        "texts": [
            "16 is represented as di di \u00bc 2 p 1\u00f0 \u00de N \u00fe c s\u00f0 \u00det, for IRD 2 p 1\u00f0 \u00de N \u00fe ct, forORD 8>< >: \u00f01\u00de where d1 is the circular arc length containing starting and ending boundaries of fault by Shao et al.16 d1 \u00bc 0:5d\u00f0 \u00de 2 0:5d Hd\u00f0 \u00de 2 rd d \u00bc J d \u00f02\u00de where d is the defect angle, d is the faulty race radius. For IRD, d \u00bc i. For ORD, d \u00bc o. Contribution of RD. For roller defect r \u00bc rt, r is the angular location of defect on roller with respect to reference. When RD strikes with inner race as denoted in Figure 1, the defect amplitude Bdri is specified as When RD strikes with OR as shown in Figure 2, the defect amplitude Bdro is specified as Bdro \u00bc r r cos J 2r sin d J mod r, 2 \u00f0 \u00de \u00f0 \u00de , 04mod r, 2 \u00f0 \u00de 5 d1 r r cos J 2r , d14mod r, 2 \u00f0 \u00de 5 d d1 r r cos J 2r sin d J mod r, 2 \u00f0 \u00de \u00f0 \u00de , d d14mod r, 2 \u00f0 \u00de 5 d 8>< >: \u00f04\u00de Bdri \u00bc r r cos J 2r : sin d J mod r, 2 \u00f0 \u00de\u00f0 \u00de , 04mod r, 2 \u00f0 \u00de5 d1 r r cos J 2r , d14mod r, 2 \u00f0 \u00de5 d d1 r r cos J 2r : sin d J mod r, 2 \u00f0 \u00de\u00f0 \u00de , d d14mod r, 2 \u00f0 \u00de5 d 8>< >: \u00f03\u00de where peak additional deflection Xr is described by Patel and Upadhyay18 for roller bearing as r \u00bc r r cos J 2r \u00f05\u00de Contribution of IRD"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000486_1.4032993-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000486_1.4032993-Figure2-1.png",
        "caption": "Fig. 2 Picture of a disk\u2013ball system captured from experiment apparatus",
        "texts": [
            " In summary, we build an integrated model for the disk dynamics by combining the governing equation (5) with the rolling friction model Eq. (6) and Coulomb\u2019s friction law Eq. (8). The model is subjected to a normal constraint equation (7) and possible nonholonomic constraints Eq. (9) responsible for a stick state in contact motion. Once physical parameters involved in the integrated model are given and an initial state of the disk motion is specified, simulation is allowed to proceed. We conducted our experiments with a disk\u2013pendulum apparatus that was first proposed in Ref. [31]. Figure 2 shows a picture of the experimental setup: The apparatus consists of a steel pedestal, a steel ball pendulum, and a steel disk. Two pendulum lines were strung to the ball via a small metal tube welded on the top of the ball and were stiff enough to balance the centripetal force experienced by the ball as it traveled down along an arc. Initially, the ball was attracted to the electromagnet owing to the electromagnetic force, and the disk vertically stood with Euler\u2019s nutation angle h0 \u00bc p=2 on the pedestal",
            " By adjusting the hitting position, the impact could create different initial conditions for the disk motion. We measured the disk\u2019s position using a stereoscopic vision method [32]. Two high-speed cameras (Lavision HighSpeedStar 4G) were used to photograph the disk, each with a resolution of 512 512 pixels at 256 gray scales. Their orientations were carefully tuned in order to capture the best image quality. The cameras were synchronized by a video distributor at a frame rate of 1000 fps. Six dots (Fig. 2) were marked on the impacted surface to facilitate image recognition. Once the disk was impacted by the pendulum, the two synchronized cameras recorded its motion. The disk motion for all six degrees-of-freedom was then reconstructed using an in-house software that provides resolutions with an accuracy of 0.005 mm in position and 0.05 deg in rotation. Detailed illustrations for data processing have been introduced in Ref. [29]. In this section, we will present the disk\u2019s physical parameters and its initial states under different test cases, and then carry out numerical simulations via the integrated model proposed in this paper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001208_j.triboint.2019.105999-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001208_j.triboint.2019.105999-Figure10-1.png",
        "caption": "Fig. 10. Distribution of contact pressure under 5 MPa compensated pressure.",
        "texts": [
            "1, and the friction coefficient between pressure compensation seal ring and sealing groove is 0.22. During the initial interference installation, the piston rod moves to the sealing slot to simulate the pre-compression deformation of the seal. Then 4 MPa fluid pressure P0 and 5 MPa compensation pressure P1 are applied on the sealing surface exposed to the hydraulic oil medium. Finally, the reciprocating speed of piston rod is set to 50 mm/s. Looking at the simulation results, the distribution of contact pressure is shown in Fig. 10. Under the action of compensation pressure and system oil pressure, the maximum contact pressure of piston rod seal contact surface is 6.79 MPa. The compensation pressure P0 is 4 MPa, 5 MPa, 6 MPa, 7 MPa, the friction stress \u03c4\u00f0x\u00de distribution and the contact pressure P\u00f0x\u00de distribution of the rod seal pair are shown in Fig. 11. The ANSYS analysis results were extracted to obtain the maximum contact pressure Pmax, maximum friction stress \u03c4  max and length L of the piston rod sealing contact surface sealing contact area under different compensation pressures, as shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003922_28.806046-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003922_28.806046-Figure10-1.png",
        "caption": "Fig. 10. Self-inductance seen by the rotor current distributions, saturated motor.",
        "texts": [
            " When the self-inductance of a stator winding was examined, a periodic variation due to rotor slotting was observed. However, there is no such variation due to stator slotting present in the self-inductance seen by a rotor current distribution. The stator slotting produces a current distribution in the rotor which has a different spatial variation from that of the fundamental -axis and -axis current distributions, and so has no effect on the self-inductance seen by these distributions. The effect of saturation on the self-inductance seen by a rotor current distribution is clearly visible in Fig. 10. There is a periodic variation in inductance at a frequency of 2 Hz about a mean value. This variation at twice slip frequency is due to the rotation of the axis of saturation in the rotor. There is a 180 phase difference between the variation of and because the -axis and -axis current distributions are orthogonal. Now that the effect of saturation on the inductances , , , and is understood, the effect of skew on these inductances will be examined. Williamson, Flack, and Volschenk have demonstrated that the time-stepping finite-element model described in Section II can be extended to include skew [5]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001155_icphm.2019.8819442-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001155_icphm.2019.8819442-Figure4-1.png",
        "caption": "Fig. 4. The architecture of a LSTM memory cell.",
        "texts": [
            " Simnorm = Sim \u2212 Min(Sim) Max(Sim)\u2212 Min(Sim) (5) As can be seen from the middle plot of Fig. 3, the similarity index starts at high values around one and suddenly decreases when a fault is formed in the bearing. A threshold of 0.9 is set for the similarity index to identify the time point when bearing degradation starts. Based on this similarity index and the true RUL of the system, the HI of the bearing unit can be constructed as shown in the bottom plot of Fig. 3. As mentioned before, a LSTM model, shown in Fig. 4, is used for predicting the bearing RUL based on a selected subset of the features extracted from the vibration signals. In Subsections II-C and II-D, we have described the procedure of selecting the most informative features (input of LSTM) and constructing the HI (output of LSTM). In this subsection, a detailed description of the LSTM architecture is provided. LSTM was originally created to overcome the problem of gradient vanishing or exploding. Instead of using hidden neurons, a LSTM model uses memory cells to build the hidden layer. There is an internal state in the memory cell that connects each cell, Ct, to itself by a fixed weight. In addition to the input node, a forget gate and an output gate are added into the memory cell as shown in Fig. 4. The parameters of a LSTM model can be calculated as follows: int = \u03a6(WinxXt +Winhht\u22121 + bin) (6) it = \u03c3(WixXt +Wihht\u22121 + bi) (7) ft = \u03c3(WfxXt +Wfght\u22121 + bf ) (8) ot = \u03c3(WoxXt +Whht\u22121 + bo) (9) Ct = int \u2297 it + Ct\u22121 \u2297 ft (10) ht = \u03a6(Ct)\u2297 ogt (11) 5 where ht is the output values from the hidden layer at time t, int, it and ot are the output values of the input node, the input gate, the forget gate and the output gate, respectively, Ct means the internal state at time t, winx, wix, wfx and wox are the weights between the input layer xt and the hidden layer ht at time t, respectively, winh, wih, wfh and woh are the weights of the hidden layers between time t and time t \u2212 1, respectively, bin, bi, bf , bo are the biases of the input node, input gate, forget gate and output gate, respectively, and \u2297 is the pointwise multiplication operator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001868_med48518.2020.9183031-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001868_med48518.2020.9183031-Figure4-1.png",
        "caption": "Fig. 4. The tilt-rotor VTOL aircraft from WingCopter has four tilting propellers where the propellers on each side are always tilted together. The vehicle has been equipped with a pitot tube, a GPS receiver, a telemetry transmitter and a Pixhawk autopilot inside the fuselage.",
        "texts": [
            " (6) 5) Attitude Controller: The inner-loop attitude controller gets an attitude setpoint \u03a8sp as an input and needs to output a torque \u03c4 such that the vehicle precisely tracks the attitude setpoints. This is achieved with a cascaded architecture, where the outer loop is a P controller that generates an attitude-rate setpoint and the inner PID loop tracks this attitude rate. First, the experimental setup is briefly described from a hard and software point of view. Then, the results from the outdoor experiments are presented in detail. The remote controlled tilt-quadrotor VTOL aircraft used in this project is shown in fig. 4 and the main characteristics are summarized in tab. I. It is based on a WingCopter v1 and has been equipped with a Pixhawk autopilot, a GPS receiver, a telemetry transmitter and a pitot tube. The Pixhawk autopilot with the PX4 firmware is a common flight controller among researchers and hobbyists because it is open-source and its source code can be modified. The FPID control architecture described above was implemented in the Pixhawk firmware by modifying the modules mc att control, mc pos control, fw pos control and vtol att control"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001692_012041-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001692_012041-Figure9-1.png",
        "caption": "Figure 9. Approximation of the head of the scarab beetle (Scarabaeus) and the model of the working body of the stubble cultivator: a - projection of the head in front; b - working body of the stubble cultivator, top view.",
        "texts": [
            " Figure 8 (c) shows the graphical dependences of the traction resistance of the model of the working bodies of the spring equalizer of the soil on the speed of movement in comparison with the serial model. An experimental model of the working bodies of the spring soil leveler provides a decrease in traction resistance by 7 ... 8% and increases the stability of the course in depth 1.54 times. The lumpiness of the soil layer decreases 2.3 ... 2.4 times, and the loosening ability is higher by 23 ... 25% in the model of the working bodies of the spring soil leveler, in comparison with ESM-5.6A. As shown in Figure 9 (a), the biological structure of the head of the prototype scarab beetle (Scarabaeus) by the nature of vital activity is of significant interest for the bionic substantiation of the working bodies of the stubble cultivator. TSIA 2019 IOP Conf. Series: Earth and Environmental Science 488 (2020) 012041 IOP Publishing doi:10.1088/1755-1315/488/1/012041 One of the main functions of loosening in a scarab beetle (Scarabaeus) is performed by the fan shape of the front of the head. The fan-shaped structure of the head of the scarab beetle shows the presence of two pairs of burrowing teeth: central and lateral. Burrowing teeth are located symmetrically with respect to the longitudinal axis. The geometry of the structure of the teeth is characterized by three indicators: the width of the teeth 1b and 2b their spacing factors 1k and 2k , with certain steps of the spacing 1S and 2S , as can be seen from Figure 9. Measurement of the geometry of digging teeth showed that their placement coefficients for the central 1k and lateral 2k , respectively, are 0.3 and 0.22. Moreover, the ratio of the width of the central teeth to the extreme will be 17.1 1 2 == b b k\u0437 . The obtained geometric features of the physical structure of the head allow us to adapt its elements to the working bodies of the stubble cultivator. The main structural elements of the model of the working bodies of the stubble cultivator are a C-shaped rack 1 with a chisel 2 and two wings 3, shown in Figure 10 (a). In figure 9 (b) the bit 2 is made expanding from the bottom up and has in its active part a cutting edge 4 with four longitudinal protrusions 5 and three longitudinal hollows 6 of a rounded shape. Two wings 3 have cutting edges 7, which are made in the form of four teeth 8 and three depressions 9, as shown in Figure 9 (b) made in the form of a sinusoid [15]. Reducing the energy costs of cutting the soil layer provides a serrated shape at the surface of the cutting edge of the bit of the working bodies of the stubble cultivator, where the tops of the teeth become stress concentrators and, with much less pressing force, cause processes of soil deformation and destruction. When designing, it is necessary to take into account that the teeth on the working surface of the cultivator bit are located along its cutting edge, which is an arc of a circle of length and radius, as can be seen from Figure 9 (b). The length of the arc of a circle can be represented by an equation of the form: ,2 121 bSSl ++= (1) where 1S and 2S are the tooth pitch. Equation (1) makes it possible to determine the number and width of the teeth depending on the length of the cutting edge of the bit. The length of the arc of the circumference of the cutting edge of the bit is described by the curved shape of the eccentric circle, which is characterized by the constancy of the angle between the normal and the radius vector, and the value of this angle is equal to the angle of internal friction of the soil, which allows the soil and plant residues to slip with minimal energy consumption"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001526_s00170-020-04987-7-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001526_s00170-020-04987-7-Figure5-1.png",
        "caption": "Fig. 5 Case study parts with different support structure configurations (artifacts 1, 2, and 3)",
        "texts": [
            " The part was printed with three different support structure configurations. The first configuration was limited to supports at the base of the part and inside two holes at its top (artifact 1), the second configuration included support structures for all overhang areas having an angle below 45\u00b0 in respect to the build plate (block, gusset, and line support structures were used in the latter case), while the third configuration was identical to the first, aside a fulldensity disk as pedestal at the base (Fig. 5). Note that the main objective of using artifacts 1 and 2 was to assess the impact of the different support structure configurations on calculated and experimentally measured distortions. The objective of printing artifact 3, which differs from artifact 1 only by the presence of a disk at its base, was to decrease the risk of part distortions when removed from the building plate. This removal was necessary for microcomputed tomography observations (see Section 2.6.6.2 for explanations). Distortions of the case study parts were computed for the 1st and 2nd support configurations using the best combination of the strain and stress modes identified from the previous study (Section 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002977_s00521-021-05986-9-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002977_s00521-021-05986-9-Figure1-1.png",
        "caption": "Fig. 1 IPC system",
        "texts": [
            " Since the dimension of the feedforward function is the same as that of the control input, the computational efficiency is improved. Moreover, the adaptive friction compensator is easy to implement since it does not need to establish detailed friction model. Experimental results demonstrate that our control strategy can lead to much more satisfactory tracking performance compared with the existing NN control strategy. Keywords Discrete-time nonlinear output regulation Inverted pendulum on a cart Tracking control Neural network Adaptive friction compensation Figure 1 shows the inverted pendulum on a cart (IPC) system, where d and Mc are the position and mass of the cart, respectively, b 2 \u00f0 p; p\u00de, Mp, Lp and Ip are the angle, mass, 1/2 length, and moment of inertia of the pendulum, respectively, u is the control force, and g is the gravitational constant. With \u00bdx1; x2; x3; x4 > \u00bc \u00bdd; _d; b; _b >, the discrete-time model of the IPC system without friction is described as follows [1]: & Yunzhi Huang hqyz@hfut.edu.cn 1 School of Electrical Engineering and Automation, Hefei University of Technology, Hefei 230009, China 2 Anhui Engineering Technology Research Center of Industrial Automation, Hefei 230009, China 3 Department of Automation, Shanghai Jiao Tong University, Shanghai 200240, China x1\u00f0k \u00fe 1\u00de \u00bcx1\u00f0k\u00de \u00fe Tx2\u00f0k\u00de x2\u00f0k \u00fe 1\u00de \u00bcx2\u00f0k\u00de \u00fe T Mp \u00feMc L2pM 2 p\u00f0cos x3\u00f0k\u00de\u00de 2 Ip\u00feL2pMp u\u00f0k\u00de \u00fe LpMpx 2 4\u00f0k\u00de sin x3\u00f0k\u00de L2pM 2 pg sin x3\u00f0k\u00de cos x3\u00f0k\u00de Ip \u00fe L2pMp x3\u00f0k \u00fe 1\u00de \u00bcx3\u00f0k\u00de \u00fe Tx4\u00f0k\u00de x4\u00f0k \u00fe 1\u00de \u00bcx4\u00f0k\u00de \u00fe T LpMc \u00fe LpMp\u00f0sin x3\u00f0k\u00de\u00de2 \u00fe Ip\u00f0Mp\u00feMc\u00de LpMp \u00f0Mp \u00feMc\u00deg sin x3\u00f0k\u00de u\u00f0k\u00de cos x3\u00f0k\u00de LpMpx 2 4\u00f0k\u00de sin x3\u00f0k\u00de cos x3\u00f0k\u00de y\u00f0k\u00de \u00bcx1\u00f0k\u00de \u00f01\u00de where T is the sampling time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002559_j.mechmachtheory.2020.104218-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002559_j.mechmachtheory.2020.104218-Figure2-1.png",
        "caption": "Fig. 2. Illustration of parameterized coordinates for a spherical triangle with side lengths of s 1 , s 2 , and \u03b1. In this case, the position vector of point C is expressed in terms of position vectors of points A and B .",
        "texts": [
            " With parameterized coordinates, the position vector of a third point on the link can be obtained as a linear combination of three vectors, i.e., the position vectors of the other two points and their product. A similar formulation was earlier introduced and used for planar linkage synthesis [23] , whereby a unified approach of linkage synthesis applicable to all three Stephenson linkages was developed. In this work, we extend the parameterized coordinates to spherical mechanisms. We take an arbitrary spherical triangle ABC, shown in Fig. 2 , as an example to describe the parameterized coordinates. It is well known that three non-coplanar vectors consist of a basis of R 3 . With reference to Fig. 2 , the position vector of point , noted by the unit vector n 3 , can be expressed in terms of the coordinates of points A and B, assuming they are known, as n 3 = \u03bcn 1 + \u03bdn 2 + \u03bbn d (3) with n d = n 1 \u00d7 n 2 (4) where n 1 and n 2 are unit vectors parallel to OA and OB . Moreover, \u03bc, \u03bd and \u03bb are dimensionless parameters, which are the parameterized coordinates of point C. They can be uniquely determined from Eq. (3) as described presently. We premultiply vectors n T 1 , n T 2 , and n T 3 separately on both sides of Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000623_978-3-319-33714-2_32-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000623_978-3-319-33714-2_32-Figure1-1.png",
        "caption": "Fig. 1 The Lucy robot. The diagram on the left shows the body segments, the position of the vestibular system, the joints and the 14 DOFs. Every joint is equipped with encoders of joint angle and angular velocity proprioceptive input) and torque. The middle shows a picture of Lucy. Actuation was force controlled. The right gives the conventions used",
        "texts": [
            " Humanoid balancing research is mostly focused on the solution of the control problem rather than on the recording, analysis, and modeling of movements. Also, humanoid balancing is often based on the zero moment point control or related measures, which try to keep the center of pressure within the base of support under the feet (Goswami 1999). For controlling biped stance without making assumptions about the support surface, the robot profits from the use an inertial measuring unit (IMU). In this work the bio-derived vestibular sensor presented in (Mergner et al. 2009) is used (Fig. 1). The DEC model is based on studies of human posture control and movement perception. First developed for the control of a single inverted pendulum body model (Mergner et al. 2006), its application has been extended to a double inverted pendulum (DIP), including hip and ankle joints (Hettich et al. 2014). Furthermore, the DEC control also was generalized for multiple DOFs and tested in simulations (Lippi et al. 2013) and in robotic experiments (Zebenay et al. 2015). This generalization was implemented in terms of a modular control architecture, in which for each DOF a control module is controlling a specific target variable",
            " The computation of the COM is based on lower limbs position, assuming that the mass distribution is known. The desired torque for the virtual joint is distributed on the four actuated joints (ankles and hips). The ankle joints are producing only a small part of the torque (\u223c20 %). The reason is that due to the relatively short foot length the foot-ground contact may easily be lost in the frontal plane. The control system applied to the frontal plane is shown in Fig. 2. The control system has been tested on the Lucy robotic platform (Fig. 1). The robots construction was inspired by human-like anthropometrics (Winter 2009). The estimators in the controls had time delays of 60 ms. Experiments consisted of performing voluntary lateral trunk lean movements (smoothed ramp-like trajectory, 4 o ) and of applying support surface tilts in the frontal plane (see Fig. 3). The robot was controlled with the DEC concept in the frontal plane as described above and in the sagittal plane as described before in the Posturob II robot (see (Hettich et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002357_j.triboint.2020.106481-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002357_j.triboint.2020.106481-Figure3-1.png",
        "caption": "Fig. 3. Elastic Deformation due to squeezing motion and oil film thickness.",
        "texts": [
            " From parametric analysis of steady state EHL equations solutions, dimensionless central film thickness is given by Dowson and Hamrock et al. [11,36] as a function of dimensionless load, dimensionless speed, dimensionless material property and ellipticity ratio \ud835\udc3b = \ud835\udc3b(\ud835\udc4a ,\ud835\udc48,\ud835\udc3a, \ud835\udc58) as in Eq. (14) [38]. \ud835\udc3b\ud835\udc50 = 2.69\ud835\udc480.67\ud835\udc3a0.53\ud835\udc4a \u22120.067(1 \u2212 0.61\ud835\udc52\u22120.73\ud835\udf05 ) (14) 2.2.2. Elastic deformation For the inner ring center is radially displaced by \ud835\udeff for outer ring is assumed to be fixed, the total squeezing of a ball along the normal of the ball-ring contact center (\ud835\udc650, \ud835\udc660) can easily be calculated from Fig. 3. In this case, elastic deformation along the ball-ring contact angle direction; can be calculated as the sum of film thickness and mutual approach by Eq. (15), from the film thickness equation given in Eq. (11) [39]. (\ud835\udeff1\ud835\udc57 + \ud835\udeff 2 \ud835\udc57 ) = (\ud835\udc51\ud835\udc56,\ud835\udc57 + \ud835\udc51\ud835\udc5c,\ud835\udc57 ) \u2212 (\u210e\ud835\udc56,\ud835\udc57 + \u210e\ud835\udc5c,\ud835\udc57 ) (15) Hereby, each ball contact in the bearing can be modeled as a Kelvin\u2013Voigt element shown in Fig. 4 and the contact force can be determined iteratively for a given mutual approach by the quasi-static method [40,41]. 2.2.3. EHL Contact force The EHL contact force for the mutual approach \ud835\udeff, squeezing speed "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000816_s1061830916120020-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000816_s1061830916120020-Figure1-1.png",
        "caption": "Fig. 1. Sample drawing that illustrates f law sizes, view from above.",
        "texts": [
            " The installation is equipped with an Yb:YAG (ytterbium-doped yttrium aluminum garnet) fiber-optic diode-pumped laser that has the maximum power of 400 W and operates at a wavelength of 1069 nm [10]. All in all, three rectangular-parallelepiped samples were manufactured by layer-by-layer laser melting. Artificial internal defects were introduced in the samples and imitated incomplete powder fusions with dimensions from 0.2 \u00d7 0.2 \u00d7 0.2 to 0.5 \u00d7 0.5 \u00d7 1.0 mm and thicknesses of 15.6 and 16.0 mm. Ten flaws with openings of 0.2 and 0.5 mm and depths from 5 to 10 mm were created in the samples. An example of the sample drawing that illustrates the sizes of the f laws is presented in Fig. 1. The f laws were created in the following manner. Domains with the above sizes were represented by cavities in the design 3D model of a sample that was fed into the installation. These domains were \u201cskipped\u201d (not fused) during the sample synthesis, whereas subsequent layers above the f laws were melted in the normal mode. Incomplete-fusion domains with prescribed dimensions were thus created in the synthesized sample, although their shapes and sizes could have differed from the ones assumed in the 3D model due to the fact that several powder layers are melted through during the synthesis, with the top layers of a cavity being fused to a greater degree than in the model",
            " The results of studying the above-described samples were represented as C-scan images. Some of the results for the most representative sample are shown in Figs. 2\u20135. This sample is peculiar in a sense that two natural f laws (the echo-signals from those are denoted in the figures as D1 and D2, respectively) were revealed in addition to the artificial f laws (the echo-signals from those are denoted in the figures by digits 1\u20134). When scanning, the samples were positioned similar to what is shown in Fig. 1, i.e., the artificial f laws were situated so that the size was increasing from left to right. Color gradation in the images that are shown in the figures corresponds to a range of amplitudes from 0 to 100% of the defectoscope\u2019s screen. It was established that with a PET with central frequency 10 MHz, a f law with the size of 0.2 \u00d7 0.4 mm could be confidently revealed at depths from 5 to 10 mm with a signal-to-noise ratio of more than 3. When using a PET with central frequency 5 MHz, 0.2 \u00d7 0.4-mm flaws could be revealed with a signal-to-noise ratio of more than 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001357_j.oceaneng.2019.106812-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001357_j.oceaneng.2019.106812-Figure6-1.png",
        "caption": "Fig. 6. Geometrical feature of cross section, welded joint configuration and measured points.",
        "texts": [
            " In detail, heat flux matrix during thermal analysis and mechanical stiffness matrix during mechanical analysis will be divided into sub-matrices, and arranged for different thread for parallel computation. Thus, much more threads were employed for one computation, and computing time can be significantly reduced. The main dimension of actual cylindrical leg structure is 36,000 mm in the length, 520t in weight as shown in Fig. 5, while the locations of measured sections were also illustrated. Thickness of cylinder is 110 mm\u2013115mm and the material is grade of EH36. The dimensional feature of cross section is illustrated in Fig. 6, in which the outer and inner diameters are 4,200 mm and 3,980 mm, respectively. Fig. 6 also shows the geometrical feature of welded joint. In detail, there are four racks on the outer surface of cylinder as symmetrical distribution with 17\ufffd away from the upright center line. As shown in Fig. 7 and Fig. 8, rack and cylinder H. Zhou et al. Ocean Engineering 196 (2020) 106812 were connected and fixed together with two 50 mm thickness plates, and these two thick plates were welded to rack beforehand with enough precision and strength, then they were welded to cylinder by means of V type welding groove and total 61 multipass welding seams symmetrically and simultaneously",
            " In detail, cross section of examined cylindrical leg structure was transformed from circle to elliptic as: the distance between upper edge and lower edge in upright direction shrink about 20 mm, and the distance between left edge and right edge in horizontal direction expand about 7 mm this kind of geometrical change caused by welding will influence the interaction between cylindrical leg structure and working platform, even the operation performance and efficiency for whole jack-up rig. In detail, Fig. 9 shows the cylinder straightness of four positions (right, left, down and up points as indicated in Fig. 6) at each measured section after welding, while the negative deformation means inward shrinkage and positive deformation means outward expansion. Currently, rigidity of cylindrical leg structure was improved with inner rib ring as shown in Fig. 10 (a) or pillar stiffener as shown in Fig. 10 (b) to avoid the welding distortion and ensure the fabrication accuracy. Although this practical approach can prevent the welding distortion as shown in Fig. 11 with inner rib ring in actual fabrication, the manufacturing cost and procedure will be increased and the total weight will be increased with about 40t when some rigidity support components cannot be removed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002234_j.matpr.2019.12.208-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002234_j.matpr.2019.12.208-Figure7-1.png",
        "caption": "Fig. 7. Modeling of composite.",
        "texts": [
            "matpr.201 in the direction of the axis of the specimen or fiber angle is 0 degree with the axis of the specimen) in present study. The five specimens are taken as per the ASTM standard and average of five is evaluated for result. Tensile testing setup is shown in Fig. 5 and flexural testing setup is shown in Fig. 6. The FEA is used for the simulation of the composite using ANSYS 15.0. For this, the modeling of composite is carried out as per the actual size of specimen of composite as shown in Fig. 7. Here, in modeling of composite, rectangular strip of composite is prepared directly instead of preparing layer by layer modeling of fiber and resin because the slipping issue faced due to smoothness of viscose rayon fiber in testing of viscose rayon fiber, as a result, the experimental elastic modulus of fiber cannot be found. The analysis of composite is carried out by assigning the properties of elastic modulus and force/load to composite specimen as fiber; (c) Viscose rayon fiber. ations of bamboo, cotton and viscose rayon fiber reinforced Unidirectional 9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003977_s0003-2670(97)00590-4-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003977_s0003-2670(97)00590-4-Figure3-1.png",
        "caption": "Fig. 3. Scheme of FIA apparatus for IEMA of cocaine.",
        "texts": [
            " In contrast, cocaine-complexed pAb\u00b1ALP conjugate is collected after passing the column. Finally, cocaine is quanti\u00aeed by a spectrophotometric or a biosensor measurement of the ALP activity where the resulting signal is proportional to the cocaine concentration in the sample. The \u00afow injection assay format separates analytebound and analyte-free conjugate easily and fast. A simple FIA apparatus consisting of a peristaltic pump, an injection valve, a column containing the immunosorbent and a fraction collector was assembled and used in the IEMA for cocaine (Fig. 3). The af\u00aenity of the antibody to its analyte determines the sensitivity of the immunoassay. The af\u00aenities of polyclonal sheep anti-BZE antibody towards related compounds were measured by BIAcore (data from Boehringer Mannheim). The apparent af\u00aenity of the IgG for BZE is 3.9 109 l mol\u00ff1. The resulting cross reactivities relative to cocaine (100%) were 12.5% for BZE and 0.1% for ecgonin. According to these data, a high degree of interaction of the antibody with cocaine in solution and with immunosorbent can be expected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure7-1.png",
        "caption": "Fig. 7. Twists diagram of 3-R(RRR)R+R HAM.",
        "texts": [
            " (18) can be written as follows: \u25e6Gi Qi = \u25e6r + \u25e6RC C Ci Qi \u2212 \u25e6Pi \u2212 \u25e6Pi Gi (18) The position vector of the moving coordinate system is \u25e6r = [r sin \u03ba cos \u03c2, r sin \u03ba sin \u03c2, r cos \u03ba]T , and the attitude matrix \u25e6RC of the moving coordinate system is as follows: \u25e6 RC = \u239b \u239c\u239d s2\u03c2\u03bd\u03b4 + c\u03b4 \u2212s\u03c2c\u03c2\u03bd\u03b4 c\u03c2s\u03b4 \u2212s\u03c2c\u03c2\u03bd\u03b4 c2\u03c2\u03bd\u03b4 + c\u03b4 s\u03c2s\u03b4 \u2212c\u03c2s\u03b4 \u2212s\u03c2s\u03b4 c\u03b4 \u239e \u239f\u23a0 (19) where s\u03c2 = sin \u03c2 , c\u03c2 = cos \u03c2 , \u03bd\u03b4 = 1 \u2212 cos \u03b4. According to the length constraint condition \u2223\u2223\u25e6Gi Qi \u2223\u2223 = L of the link, the following relationship can be obtained. ( xGi \u2212 xQi )2 + ( yGi \u2212 yQi )2 + ( zGi \u2212 zQi )2 = L2 (20) Therefore, we can get the actuation angles \u03b8i1(i = 1, 2, 3) of 3-R(RRR)R+R HAM. As shown in Fig. 7, when the 3-R(RRR)R+R HAM moves in a certain range according to pitch or azimuth motion. Each revolute joint of the HAM rotates around its axis. \u03c9i j (i = 1, 2, 3; j = 1, 2, \u00b7 \u00b7 \u00b7 5) and \u03b8i j , respectively, represent the axis vector and rotation angle of the i-th revolute joint of the j-th branch chain. \u03c96 and \u03b86, respectively, represent the axis vector and rotation angle of the revolute joint of the polarization mechanism. The motion subspace of revolute joint is expressed as \u03be\u0302r = [0, 0, 0, 0, 0, 1]T (21) https://doi"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001166_j.procir.2019.03.197-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001166_j.procir.2019.03.197-Figure5-1.png",
        "caption": "Fig. 5. Parameterized CAD model of 3RRR manipulator in initial configuration.",
        "texts": [
            " Summing up, the CAD system enables the adjustment of the model as well as the export for the examination of the inequality conditions while the slicing software calculates the objective. 4. Application In the following, the presented method is applied to a case study. By choosing the printer Ultimaker S5, it is possible to print support structures with water-soluble filament with an additional second extruder and thus enables the AM of nonassembly mechanisms. 4.1. Presentation of the case study In order to show the applicability and the benefits of the optimization method, a planar 3RRR manipulator is used as a case study. According to Fig. 5 three arms are attached to the triangular base plate of the mechanism, each consisting of two links. By fixing the base plate, six rotational DOFs are unconstrained and thus have to be parameterized in the CAD model. Due to the layerwise build process, the geometric accuracy of the guiding Martin Hallmann et al. / Procedia CIRP 84 (2019) 271\u2013276 275 M. Hallmann et al. / Procedia CIRP 00 (2019) 000\u2013000 4 to the predefined conditions. In a final step, the toolpaths for the chosen, optimal solution can directly be used for the AM",
            " Procedure of CAD-integrated build time and support material quantity optimization using slicing software for detailed-analysis-based build time estimation. 4. Application In the following, the presented method is applied to a case study. By choosing the printer Ultimaker S5, it is possible to print support structures with water-soluble filament with an additional second extruder and thus enables the AM of nonassembly mechanisms. 4.1. Presentation of the case study In order to show the applicability and the benefits of the optimization method, a planar 3RRR manipulator is used as a case study. According to Fig. 5 three arms are attached to the triangular base plate of the mechanism, each consisting of two links. By fixing the base plate, six rotational DOFs are unconstrained and thus have to be parameterized in the CAD model. Due to the layerwise build process, the geometric accuracy of the guiding 4 M. Hallmann et al. / Procedia CIRP 00 (2019) 000\u2013000 5 surfaces of the revolute joints are influenced by their orientation. To avoid staircase effects, the mechanism is positioned flat on the platform so that the axes of the joints are equal to the build direction z (see Fig. 5). 4.2. Optimization As it is known from section 3, any stochastic optimizer, which can handle nonlinear constraints, is suitable for the build time and support material quantity optimization. In this contribution, a Genetic Algorithm is examplarily applied to the given case study. Although metaheuristic algorithms are problemindependently applicable, the results are strongly dependent on the settings of the optimizer, which have to be chosen with respect to the given optimization problem. Based on previous studies, the population size was set to 50, the number of generations to 200"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure12-1.png",
        "caption": "Fig. 12. Scoop-Tube Coupling with Hagan Displacement Control Cylinder",
        "texts": [
            " 1, a very high standard has been achieved by using a simple reversible gear pump which is normally idle, and is Elimination of Auxiliaries. operated only during the moments when the speed is being adjusted. I t is, however, in the case of automatically controlled drives requiring a continuously running servo-pump and piston valve for filling or emptythat the question of dependence upon auxiliaries is a reasonable point of criticism. An improvement in this direction is the displacement cylinder control evolved in conjunction with the Hagan automatic regulator, as illustrated by Fig. 12. The coupling is of the standard scoop tube type, the reversible motordriven pump and control switch being retained as a complete standby for remote manual control. The automatic control cylinder is con- 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 97 nected by pipe to the scoop tube discharge branch, its piston being actuated by the compressed air or other fluid pressure used for the remainder of the combustion control system. The quantity in the working circuit can be quickly increased by raising the control piston, and similarly decreased by lowering the piston"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003720_s0039-9140(96)02040-1-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003720_s0039-9140(96)02040-1-Figure10-1.png",
        "caption": "Fig. 10. Microcell with a working rotating disk electrode.",
        "texts": [
            " In the case of the vibration pump [81], the fall of the mercury drop used as the working static merucry electrode was eliminated. As the rotating working electrode, a disk electrode is usually used [84]. The rotating disk in microcells has been made from the one of following materials: glassy carbon without [5] and with a mercury film [28], Au [20], Pt [20,21] and impregnated graphite with a mercury film [36,37]. The sample volume is 0.5-4.0 ml [20,21,23] but the sample volume can be also decreased to 200/~1 [28,36,37]. A microcell with a rotating disk working electrode is shown in Fig. 10 [28]: (1) is the rotating disk working electrode, (2) is the auxiliary electrode, (3) is the reference electrode and (4) is the sample. The auxiliary and reference electrodes are separated by electrolytic plugs (5) of porous ceramics. It should be noted that forced convection has been successfully combined with hydrodynamic modulation [85-93]. Although this effect was applied in macrocells, it can also be used in microcells to decrease the influence of the current which is not connected with the diffusion of the analyte to the working electrode surface"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001315_tiv.2019.2955904-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001315_tiv.2019.2955904-Figure12-1.png",
        "caption": "Fig. 12. Tip-over constraint: (a) Clockwise tip-over, and (b) counterclockwise tip-over.",
        "texts": [
            " The reaction force R of the vehicle in (29) must be nonnegative, to maintain contact with the ground, which yields s\u0307 \u2264 \u221a \u2212 gkN \u03banN . (35) Because kN , g, and \u03ba are always positive, Eqn. (35) only applies to cases where nN is negative. In other words, the angle between N and n must be greater than or equal to 90\u25e6. When the entire weight of the vehicle shifts to the left/right side of the vehicle and the tires on the right/left are about to lose contact with the ground, the vehicle is about to tip over. The condition for not tipping over in the counterclockwise direction is fq \u2265 \u2212R b h\u0302 , (36) where h\u0302 and b are defined in Fig. 12. If we denote \u03b7 = b h\u0302 , then the condition for the clockwise direction is fq \u2264 R\u03b7 . (37) Combining (36) and (37) into one constraint equation yields f2q \u2264 (R\u03b7) 2 . (38) Substituting (28) and (29) into (38) leads us to a quadratic inequality in s\u0308 and s. The solution of the quadratic inequality (A-5) derived by [7] is given in the appendix. In this study, the angle between t and the vector along the longitudinal axis of the vehicle \u03b8 is zero, as the planar path of the vehicle is determined from the position (x, y) of the midpoint of the rear axle, and hence the tip-over constraint does not impose any constraints on the acceleration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001096_s00170-019-04076-4-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001096_s00170-019-04076-4-Figure14-1.png",
        "caption": "Fig. 14 Grid of tooth surface of worm wheel",
        "texts": [
            " These deviations were small because the deviations do not seem to have an influence on the tooth contact patterns. The tooth surfaces of the machined worm wheel were measured using a coordinate measuring machine (Gleason Works Sigma M&M 3000) and were compared with the nominal data on the tooth surface that were determined mentioned above. In this case, five points in the direction of the tooth profile and nine points in the direction of the tooth trace for the grid were used. The reference point is defined in the center of the tooth surface and the deviation is zero at this point. Figure 14 shows the grid of the tooth surface of the worm wheel in the case of R = (x2 + y2)1/2. Figure 15 shows the measured results of the tooth surface deviations of the worm wheel. The maximum values of the magnitude of deviations are 0.035 mm and 0.025 mm on the left and right flanks, respectively. This value as a whole is a permitted limit enough because the reason is that the size of the worm wheel is very large and seems not to have an influence on the tooth contact pattern. The surface roughness of the machined worm wheel was measured using a surface roughness measuring instrument (Mitutoyo SJ-210)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003134_s11012-021-01388-2-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003134_s11012-021-01388-2-Figure1-1.png",
        "caption": "Fig. 1 Waterbomb pattern and natural way of folding. a 6 creased waterbomb unit cell with sizes 2a and 2b and inner angle k, b illustration of the geometric relationship between natural force and waterbomb fold pattern, c illustration of a closed m n waterbomb tessellation, with m \u00bc 3 lines and n \u00bc 8 cells in each line; d Illustration of a closed m n waterbomb tessellation, with m \u00bc 6 lines and n \u00bc 20 cells in each line",
        "texts": [
            " In this regard, it is possible to imagine a total symmetry where the fundamental substructure is the unit cell, or other situations where the fundamental structure is built by a set of unit cells, RVE. Symmetry conditions might need to be associated with both geometrical and mechanical aspects. The complexity of the origami description increases as the asymmetry of the fundamental representative element accentuates. Therefore, the complexity of the closed tessellation representation relies mostly on the unit cell configuration, regarding its degree of asymmetry. This paper considers a 6 creased waterbomb pattern (Fig. 1a), which closed tessellation results in a cylindrical-like structure. The natural force associated with the folding process of the closed tessellation tends to generate a conic-like structure (Fig. 1b) [15, 24], which results in local asymmetries even for a symmetric actuation. This conic-like natural force can generate several 3D structures, including a ballshaped origami (Fig. 1c) and a cylindrical origami (Fig. 1d). The following sections present a local asymmetry study, evaluating the behavior of a unit cell of the waterbomb pattern considering both kinematics and mechanical approaches. The analysis of the waterbomb pattern can be based on the hypothesis of rigid origami, where all the deformation is localized on the creases (mountain and valley folds), which means that faces remain flat and undeformed. Therefore, it is possible to analyze the waterbomb unit cell as a mechanism (Fig. 2), where the creases are represented by revolute (or cylindrical) joints and the faces are represented by rigid links",
            " The points in the workspace should not be considered individually, but as a set of 6 points that define the unit cell. As examples, 9 sets are selected where each one corresponds to a subset of angles h1; h2; h3\u00f0 \u00de given as an input and a subset of angles h4; h5; h6\u00f0 \u00de that corresponds to the output. The subset for each case in Fig. 4a\u2013i is shown in Table 1, along with the converged remaining three angles, OF = b ffiffiffi 2 p , and vertices B and E belong to the inner sphere with radius OB = OE = b (see Fig. 1a) Table 1. Figure 4a shows the particular case where the waterbomb is fully unfolded. A different perspective of the workspace is now of concern highlighting the vertices on the workspace. Figure 5 highlights each joint-linkage pair, considering a referential frame at creaseOA (joint A). It should be highlighted that superposition and penetration of panels define unfeasible regions associated with an empty space on the workspace. In other words, holes and isolated groups are generated in the workspace due to unfeasible configurations",
            " 4, can be extrapolated as representative of the structure behavior, and the influence of the endings is contained within the cells from the first and the last line. In addition, a tessellation with an even number of cells per line (Fig. 17b) presents a significant deviance from the symmetric behavior (caseP3). In opposition to the result presented in Fig. 17a, Fig. 17b indicates a smoother increasing of the maximum deviance. Besides, the increase of the number of cells causes the loss of the circumferential symmetry. This fact is due to the folding process that occurs on the tessellation. The natural force (Fig. 1b) associated with the folding process of the closed tessellation evaluated in Fig. 17 tends to generate a ball-shaped origami, where the middle lines tend to present a symmetric behavior, and the endings tend to be asymmetric. On the other hand, if this natural force (Fig. 1b) tends to generate a cylindrical-like origami, a different behavior is observed, where the middle lines tend to present an asymmetric behavior and the endings tend have a symmetric behavior. The folding process of a cylindrical-like origami is shown in Fig. 18. The structure presents a motion that can be translated as a circumferential reduction, followed by an axial compression, a strangulation, and an axial relaxation. This motion sequence is presented at Fig. 18a, for a tessellation with m = 5 lines, and at Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure15-1.png",
        "caption": "Fig. 15 Sketch of the grinding wheel posture for rake face grinding",
        "texts": [
            " Therefore, this paper just focuses on the grinding method of the curved rake face. To fit with the shape characteristics of the rake face, a pointgrindingmethod is adopted by use of the edge of a grindingwheel. In the grinding process, the edge of the grinding wheel contacts with the surface to remove material, and the whole surface is groundpoint by point under five-axismovements. In consideration of the uneven shape of the rake face, a single-sided bevel wheel is used in order to avoid interference or over-cutting. The posture of the grinding wheel is shown in Fig. 15. The radius of the grinding wheel is required to be less than the minimum curvature radius to avoid partial interference. According to the theory of differential geometry, the mean curvature can be calculated by: Hi LiGi\u22122MiFi \u00fe NiEi 2 EiGi\u2212F2 i \u00f031\u00de Using Gaussian curvature and mean curvature, the maximum normal curvature is obtained by: \u03bamax \u00bc Hi \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffi Hi 2\u2212Ki p \u00f032\u00de Then, the minimum curvature radius is calculated by: \u03c1min \u00bc 1 \u03bamax \u00f033\u00de The radius of the grinding wheel should be less than the minimum curvature radius, that is Rs\u2264min \u03c1minf g \u00f034\u00de The NC program for the curved rake face is generated by CAM software after path planning and post-processing"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure27-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure27-1.png",
        "caption": "Fig. 27. King Valve Traction Coupling",
        "texts": [
            "comDownloaded from PROBLEMS OF FLUID COUPLINGS 121 The principal feature of this type of coupling is the ring valve for reducing the drag torque and controlling the slip by throttling the vortex circulation of the liquid. The design was originated in an attempt to produce a \u201cdeclutchable\u201d fluid coupling, which could be used in direct combination with a conventional sliding gearbox, but the results were disappointing because at high engine speeds with the ring valve closed there is quite a considerable drag torque. A typical section is shown by Fig. 27 and a set of slip curves taken at constant torque for different degrees of opening of the ring valve appear in Fig. 28. It can be seen that the initial closing movement has little effect on the slip, whereas the effect increases rapidly as the valve approaches the closed position. A disconcerting feature of the early designs was the tendency of the ring valve to draw into the circuit automatically when the load on the coupling was increased. The reasons were not studied further at the time, since the problem of reducing drag torque had been dealt with more simply by means of the reservoir chamber and the anti-drag baffle, but it has since been established that the force which draws the ring valve into the circuit is largely influenced by the shape of the valve and the area exposed at its two ends, which are in contact respectively with the vortex ring and the liquid in the reservoir chamber"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002038_j.matdes.2020.109353-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002038_j.matdes.2020.109353-Figure3-1.png",
        "caption": "Fig. 3. Geometry and sampling location of the in-situ specimen (unit: mm).",
        "texts": [
            "13 mm and the track overlap of 0.105 mm. The powder-bed depth is set to 25 \u03bcm. To reduce the anisotropy of the mechanical properties, a cross-directional (N-layer 45\u00b0; and (N + 1) layer 135\u00b0) laser scanning strategy is used. The solution treatment and aging (STA) are conducted at 950 \u00b0C for 2 h followed by air cooling (AC), and 550 \u00b0C for 4 h followed by AC, to improve the ductility and uniformity of SLM samples. After heat treatment, the samples are machined according to the drawing of the in-situ test specimens (Fig. 3). The cross-sectional area of the test piece should be less than the ratio of the maximum load of the testing machine to the tensile limit of the material to ensure the integrity of the tensile curve. The thickness of the test piece is related to the penetrability of X-ray, which is selected according to the recommendations of SSRF. To avoid the influence of surface roughness on the tomography and test process, the specimen is polished until there are no obvious scratches on the surface. The in-situ tensile test is performed on the high-resolution SRXT at the BL13W1 of the Shanghai Synchrotron Radiation Facility (SSRF) in Shanghai, China"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000175_1350650119882834-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000175_1350650119882834-Figure1-1.png",
        "caption": "Figure 1. Schematic of partial slip texture multi-lobe journal bearings: (a) Two-axial groove; (b) two-lobe; (c) three-lobe; and (d) offset.",
        "texts": [
            " The partial slip texture geometry on bearing lobe surface relative to the orientation of external load will have significant eGects on the performance of multi-lobe journal bearings. The effect of partial slip texture configuration is considered on the major load bearing lobe surface which extends from groove inlet along flow direction. The partial slip texture configuration of multi-lobe (two-axial groove, two-lobe, three-lobe, and offset) journal bearings under vertical load with angular extent of partial slip texture region ( st), and angular extents of successive regions of land with slip ( sl) and recess ( tr) regions are illustrated in Figure 1. Flow along circumferential direction based on infinite number of cells of land with slip and recess in partial slip texture regions35 in a lobe based on narrow groove theory36 using the local pressure distribution over land (pdl) and recess (ptl) is obtained as ql \u00bc ujhl 2 hl \u00fe 2b\u00f0 \u00de hl \u00fe b\u00f0 \u00de 1 h3l 12 hl \u00fe 4b hl \u00fe b dpdl Rd \u00f01\u00de ql \u00bc uj hl \u00fe htl\u00f0 \u00de 2 1 hl \u00fe htl\u00f0 \u00de 3 12 dptl Rd \u00f02\u00de The local pressure gradients over land and recess are related to the overall pressure gradient in a lobe as dpl d \u00bc dpdl d \u00fe 1 \u00f0 \u00de dptl d \u00f03\u00de The nondimensional flow along circumferential direction through land with slip and recess is related to the overall nondimensional pressure gradient in a lobe as Ql \u00bc Hl 2 1d 2d \u00fe 1 \u00f0 \u00de 1t 2t 1 2d \u00fe 1 \u00f0 \u00de 1 2t H3 l 12 1 1 2d \u00fe 1 \u00f0 \u00de 1 2t dPl d \u00f04\u00de where 0d \u00bc 1\u00fe l, 1d \u00bc 1\u00fe 2l 1\u00fe l , 2d \u00bc 1\u00fe 4l 1\u00fe l 1t \u00bc 1\u00fe , 2t \u00bc 1\u00fe \u00f0 \u00de 3 \u00f05\u00de The nondimensional form of modified dynamic Reynolds equation for multi-lobe journal bearing with partial slip texture configuration is derived as dHl dT \u00fe d d slHl 2 plH 3 l 12 dPl d \u00bc 0 \u00f06\u00de where the coefficients are sl \u00bc 1d 2d \u00fe 1 \u00f0 \u00de 1t 2t 1 2d \u00fe 1 \u00f0 \u00de 1 2t , pl \u00bc 1 1 2d \u00fe 1 \u00f0 \u00de 1 2t \u00f07\u00de The eccentricity ratio (\"l) and attitude angle ( l) of each lobe for multi-lobe journal bearings are calculated1 based on the relationships of journal center relative to bearing and lobe centers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000281_iecon.2019.8927827-Figure20-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000281_iecon.2019.8927827-Figure20-1.png",
        "caption": "Fig. 20 \u2013 Assisted reluctance rotor with bonded magnets",
        "texts": [
            " The assisted reluctance machine (PMaSRM) improves the torque density and power factor compared to synchronous reluctance machines keeping a good interval of flux-weakening. In general, the regular sintered magnets are used and inserted in flux barriers. On the other hand, recently the rotor geometries assumed more and more very complex shapes. For this reason, the flux barriers cannot be completely filled with traditional magnets. The drawback can be resolved with the use of bonded magnets [19]. For this reason, the bonded magnets have been used (PA6 NdFeB self-prepared with binder content at 10% as in Fig. 8) for assembling the assisted reluctance rotor (Fig. 20). Very promising results have been obtained concerning the torque value and the factor compared to synchronous reluctance (SRM) machine and regular anisotropic ferrite magnets assisted reluctance machine, as shown in TABLE II. The materials of this family, made by a part of hard magnetic powder and a part of soft magnetic powder, can be efficiently used as flux generators in magnetic sensor applications, thanks to their steep magnetic characteristic. The preparation of HMCs is very similar to bonded magnets by compression moulding"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure35-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure35-1.png",
        "caption": "Fig. 35. V-Belt Pulley for Auxiliary Drive with Resilient Mounting of Output Shaft",
        "texts": [
            " SELF-ALIGNING OUTBOARD BEARING I ldFdZ& / FLEXIBLE COUPLING F L ~ D COUPLING - Fig. 34. Alternative Methods of Mounting Traction Couplings 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 131 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 132 PROBLEMS OF FLUID COUPLINGS 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 133 One arrangement of V-belt drive is shown by Fig. 35 in which the pull of the V-belts is taken directly by the ball bearing mounted in the bell housing, while the overhang of the coupling is supported by a \" Silentbloc\" rubber-bushed flexible coupling which transmits only the power required by the auxiliary drive. The driven pins are mounted on the V-belt pulley and are thus carried by the outboard bearing, while the driving half is mounted on the runner shaft, which it supports 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001900_s0263574720000806-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001900_s0263574720000806-Figure16-1.png",
        "caption": "Fig. 16. The behaviour of ISS in the presence of a U-shaped object, a case where the robot is trapped in an unsafe state. The selected trajectory is depicted in green. In the first snapshots (from left), low cost collisionfree trajectories to the goal are selected. In the last snapshot (on the right), the robot is in a state where no collision-free trajectory is available. For more clarity, the robot\u2019s planning state of the previous snapshot is represented by the grey disc.",
        "texts": [
            "4, PASSPMP-PSO has been compared to PASSPMP wrt this problem. It has been found that PASSPMP-PSO has good performances to find a global optimum, while PASSPMP has been trapped in a local minimum. However, PASSPMP\u2019s safety guarantee remains valid. ISS approach has been tested in the same scenario of Fig. 14, with a U-shaped obstacle and for the same conditions and constraints. Two situations are highlighted: in the first case, the robot can end up in a situation where no collision-free trajectory is available. Figure 16 illustrates ISS results for such a situation. In the two first snapshots (on the left), a collision-free trajectory to the goal is selected (the closest to the goal). However, in the last snapshot, even if the previous trajectory drives the robot to a collision-free state, no collision-free trajectory is available from this state. Therefore, the robot is trapped in a local minimum which, in addition, is unsafe. In the second situation, the robot has found a collision-free trajectory, which has been selected closest to the goal, as illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003867_ip-cta:19951883-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003867_ip-cta:19951883-Figure1-1.png",
        "caption": "Fig. 1 Two incerred pendulums on carts",
        "texts": [
            " 16 or 22 is discontinuous with respect to its argument, which is known to display chattering phenomena [14]. As suggested in Reference 13, this problem can be avoided by the so-called boundary layer approach, that is, sign (x) is replaced by (xjnorm (x) + 6) with 6 being a small positive constant. 4 Simulation example To demonstrate the performance of the proposed decentralised adaptive controller, the numerical example used in Reference 12 will be considered here. The two inverted pendulums are connected by a moving spring mounted on two carts as shown in Fig. 1. I t is assumed that the pivot position of the moving spring is a function of time which can change along the full length I of the pendulums. The motion of the carts is specified as sinusoidal trajectories. The input to each pendulum is the torque ui applied at the pivot point. The control objective is to find suitable ui so that each pendulum tracks its own desired reference trajectory while the connected spring and carts are moving. and xz = ( i2 , I)~)', choosing k = 1 , I = 1, M = rn, the dynamic Defining the state vectors x, = ($,, 4 4 2 model for the system can be given as i = 1, 2 i i = A i x i + B i u i + Bizi where - - with y , = sin (2t) y2 = 2 + sin (3 t ) and the uncertainty u(t) = sin (5t)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.11-1.png",
        "caption": "Fig. 10.11 Blending device and working principle [2]: 1\u2014agitator; 2, 3, 4\u2014rotation system",
        "texts": [
            " The cellulose is washed at 15\u201325 \u00b0C for about 30 min. After acidification, the cellulose is washed with hot water with a temperature of 50\u201360 \u00b0C for one time and then washed with cold water until neutral. (3) Blending The blending of refined cellulose is mixing several small batches with different quality through the same mechanical stirring to form a large batch with a homogeneous quality. The blending device is a trough made of fiberglass composite or engineering plastic. The blending device for the refined cellulose is shown in Fig. 10.11. The semicircular bottom of the blending device is made of epoxy resin material. When blending to prepare a large batch, it is necessary to ensure (usually a loading amount of degreasing pot) the homogeneous quality of each small batch. Do not blend small batches with significantly different viscosities, which can minimize heterogeneous quality caused by the production process. The batch combined by blending is called a large batch or total batch. The viscosity of total batch of refined cellulose is calculated by the weighted average method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.39-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.39-1.png",
        "caption": "Fig. 10.39 The structure of disk mill: 1\u2014adjustment wheel, 2\u2014feed port, 3\u2014fixed disk, 4\u2014transmission turbine, 5\u2014rotating grinding, 6\u2014discharge port, and 7\u2014drive shaft",
        "texts": [
            " There are four sets of stator knives located evenly around the roll knife, in which the blade is parallel to the axis. Compressed air or high-pressure water is employed for adjusting the knife. The feeding inlet of the chopper is equipped with a feeding impeller, by which the NC slurry is continuously fed evenly between the rolling cutter and the bed knife. The chopped NC pulp is discharged into the second chopper. 3. Disk grinding chopper Disk mill is mainly composed of shell, spindle, fixed disk, and rotating disk. The structure is shown in Fig. 10.39. The fixed disk can adjust the distance between the disks by adjusting the device. The diameter of the rotary disk is 350 mm with an inner diameter of 100 mm. There are 104 knives (teeth) between rotation mill and fixed mill with a width of about 4 mm, space of 4 mm, and height of 10 mm. There are 8 feed tracks on the ground, which divides the disk into 8 equal parts. The depth and width of the track are 6\u20138 mm and 15 mm, respectively. Teeth and track have a certain angle. The disk material is made of 45# steel"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001602_ab87e3-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001602_ab87e3-Figure11-1.png",
        "caption": "Figure 11. Experiments of starting process of motor: (a)motor prototype under high-speed camera (b) experiment system",
        "texts": [
            " The manufactured stator is shown in figure 10(a), in which the ring stator is supported by six flexible beams located on the saddle points of B13 mode and three gaps are created between two adjacent beams to form a perfect match of two orthogonal modes. The metal layer is connected to pads through the beams and were applied by alternative voltage to deflect the PZT plates to generate traveling wave in stator for driving. The main parameters of fabricated stator are listed in table 1. The packaged motor is shown in figure 10(b). The motor prototype in black and white under high-speed camera and the whole experiment system is shown in figure 11. The rotation of motor is captured by a high-speed camera (5000FPS) and recorded by both computer and oscilloscope to analyse the starting process for the angular velocity and displacement curves of rotor. The experiment results of three different motors are shown in figure 12. 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 e t d M a us cr 6 (a) (b) (c) (d) (e) Si SiO2 PZT Bottom platinum Top platinum Figure 9"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure18-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure18-1.png",
        "caption": "Fig. 18. Pressure distribution in the axial gaps under axial static scenario and a 2000 RPM and 200 bar operating condition.",
        "texts": [
            " As the transverse contact force Fc,u acts on two gears in opposite directions, the combined transverse force experienced by two gears are different, especially in y-direction. This results in different gear eccentricity, radial sealing, and casing wear distribution on the driver gear side and driven gear side. The gear dynamics on the transverse plane (i.e. x-y plane) overlaps largely with that of the spur gear pumps, as discussed in previous works [29,38]. X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 Under the \u201caxial static\u201d scenario (i.e. vz = \u0394z = 0,htop = hbottom), the pressure distributions in the axial gaps are shown in Fig. 18. For the reference design with one circular pitch helix rotation and no relief groove, the pressure distribution in the meshing zone is inphase for the top gap and bottom gap. However, a phase shift occurs at the outer circumference (highlighted in Fig. 18), given by the helical shift of the tooth-space. This phase shift in pressure distribution results in a larger axial force from the top gap than from the bottom gap, hence a net downward axial force on the gears. For high-pressure operations, these shearing frictions are far smaller than the loads given by fluid pressure and gear contact. For simplicity, in the analysis in this subsection, the torque loss given by the shearing of the fluids in the journal bearing and axial lubrication gaps is neglected",
            " In addition, it is interesting to have a simple estimator on the mean magnitude of the axial load from the axial gear surface Fz,p2, without needing to run a computational model to solve it. Since the difference Fz,p2 = \u20d2 \u20d2Fz,p2,top \u20d2 \u20d2 \u2212 \u20d2 \u20d2Fz,p2,bottom \u20d2 \u20d2 essentially comes from the X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 pressure phase-shift at the outer circumference, it can be conjectured that the net force difference is approximately the fluid-pressure force acting on the axial surface of a full tooth in the high-pressure region (as highlighted by the \u2018pressure difference\u2019 interval in Fig. 18). Herein an estimator for the mean axial fluid-pressure force F\u0302z,p2 can be suggested: F\u0302z,p2 = cp Atooth\u0394P (60) where cp is a model coefficient. For the reference pump, the lateral surface area for a full tooth isAtooth = 62.3 mm2. From the results shown in Fig. 19, the value of cp is estimated as 0.66. With this estimate of such coefficient, one can potentially have a quick judgement for the force on the axial gear surface. Fig. 20 shows the moments experienced by the two gears in the reference pump, and the direction and magnitude of the moment components are qualitatively highlighted in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001908_j.jmrt.2020.08.101-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001908_j.jmrt.2020.08.101-Figure1-1.png",
        "caption": "Fig. 1 \u2013 Presentation of the slotting method for modeling t",
        "texts": [
            " Both cores were characterized by different densities and double cell wall sizes as shown in Table 1. It worth noting that the relative density of the core material was estimated from the proportion of the composite to the unit cell volume. For square double cell wall core structure, the relative density was calculated using Eq. (1). \u0304 = 2t(lout + liner) l2out (1) As for cores fabrication, a simple slotting technique is applied, where the flax/PLA composite panel was cut into the required dimension using a bench saw machine. Fig. 1 illustrates the slotting procedure used in fabricating the cores specimens with a double cell wall having 20 mm and 30 mm cell sizes respectively. Here, the slot area was equal to the j m a t e r r e s t e c h n o l . 2 0 2 0;9(6):12065\u201312070 12067 w t T e 2 2 W t o o p i 2 T s e p t a v t a b t W m 3 3 t T t i t p c he square interlocking cores. idth of the composite thickness. The slotting height was set o 10 mm and the rest of cores dimensions are presented in able 1. The skins were bonded to the core by applying an poxy adhesive"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002673_s40684-020-00283-7-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002673_s40684-020-00283-7-Figure5-1.png",
        "caption": "Fig. 5 The morphology of the four single-track printings in the 60\u00a0W_500\u00a0mm/s specimen",
        "texts": [
            " In order to investigate the microstructure and the geometries of single-track printing, its lateral surface was provided using a disc-type abrasive sawing machine first, and the sandpapers, #800, #2000, and #4000, were then applied to polish the surface in sequence. Electronic etching was then carried out in a solution with 10% extraction for corrosion under a voltage of 5\u00a0V for 30\u00a0s. Then, the image of the lateral surface of the printing was observed on an optical microscope (METALLUX3 Leica, Germany). Four single-pass tracks have been made for some specimens. The topographies for the four single-pass tracks of the 60\u00a0W_500\u00a0mm/s specimen are shown in Fig.\u00a05. They are presented to be a zigzag morphology such that the uniformity of asperity heights is strongly dependent on the laser operating conditions and random particle size distribution in the powder bed. The average height (H), as the red horizontal line is shown, can be evaluated for every track in the specimen. Figure\u00a06 shows one lateral surface image of the four cuttings in the 60\u00a0W_500\u00a0mm/s specimen. Height, H, depth, D and width, W, are marked. The experimental results of H, D, W for various laser conditions are shown in Table\u00a08, the ones for D and W are obtained to be the average of these two geometries formed at four cross sections of a single-pass track"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003122_s11012-021-01410-7-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003122_s11012-021-01410-7-Figure1-1.png",
        "caption": "Fig. 1 Structural diagram of a simple dual-rotor and an inter-shaft bearing system",
        "texts": [
            " Thus, the dynamics of the system is coupled with the heat transfer of the bearing. Furthermore, the viscosity-temperature relationship of the lubricant is considered in this model. Results show that the bearing\u2019s temperatures form two \u2018\u2018temperature peak\u2019\u2019 in the \u2018\u2018resonance zone\u2019\u2019 of the system, which indicates the dynamic properties of the system make an important impact on the bearing\u2019s temperatures. 2.1 Dynamic load of the inter-shaft bearing The structural diagram of a simple dual-rotor and an inter-shaft bearing system is shown in Fig. 1 [24, 25]. The LP and HP rotors are coupled by a radial cylindrical roller bearing, whose outer and inner rings whirl with HP and LP rotors. Where x1 is the LP rotor\u2019s rotation speed while x2 is the HP rotor\u2019s rotation speed. The rotation speed ratio k \u00bc x2 x1 is constant during operation, where k[ 1 for co-rotating while k\\ 1 for counter-rotating. Assume that the LP and HP rotors rotate at constant rotation speeds and the rotation speed ratio is constant, the DOFs of the rotational angles of LP and HP rotors around the z-axis can be ignored. Therefore, the simple dual-rotor system has 8DOFs, which are the vertical and horizontal displacements of LP and HP rotors x1, y1, x2, y2, and the rotational angles of LP and HP rotors around the vertical and horizontal axes hx, hy, ux, uy. The structural diagram of a simple dual-rotor and an inter-shaft bearing system is shown in Fig. 1. The second kind of Lagrange\u2019s equation is applied to establish the dynamic equations of the system. Both LP and HP rotors are assumed as rigid rotors, the energies of the system [24\u201326] are: The kinetic energy of the system is T \u00bc 1 2 m1 _x2 1 \u00fe _y2 1 \u00fe 1 2 Jd1 _h2 x \u00fe _h2 y \u00fe 1 2 Jp1 x2 1 2x1hx _hy \u00fe 1 2 m2 _x2 2 \u00fe _y2 2 \u00fe 1 2 Jd2 _u2 x \u00fe _u2 y \u00fe 1 2 Jp2 x2 2 2x2ux _uy ; \u00f01a\u00de where m1, m2 are the mass of the LP and HP rotor; Jp1, Jp2 are the polar moment of inertia of the LP and HP rotor; Jd1, Jd2 are diameter moment of inertia of the LP and HP rotor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001157_s11071-019-05220-1-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001157_s11071-019-05220-1-Figure8-1.png",
        "caption": "Fig. 8 Configuration of the single wheel set and roller rig",
        "texts": [
            " To investigate such phenomena theoretically, it is necessary to account for the quintic nonlinearity in the system by performing fifth-order nonlinear analysis. In the next section, we present a simplified equation for nonlinear hunting motion, that is a normal form, with the aid of center manifold theory and propose a method for identifying the coefficients of the normal form from the experimental results in this section. 3.1 Equation of motion First, we consider amathematicalmodel of a suspended single wheel set with two degrees of freedom, lateral motion y and yaw motion \u03c8 , as shown in Fig. 8, which has been commonly used as an essentialmodel [44\u201346]. In the figure, r0 is the centered wheel rolling radius, kx is the x direction stiffness, \u03b30 is the wheel tread angle, d0 is the half-track gauge, x is the running direction, and v is the running speed. The parameter values corresponding to the experimental apparatus used in this study are shown in Table 1. We used the half-track gauge d0 as the representative length and the inverse value of the linear natural frequency of yaw motion \u03c9\u03c8 as the representative time to obtain the dimensionless equations governing the dimensionless lateral and yaw motions y\u2217 and \u03c8 , respectively, as in [6,13]: y\u0308\u2217 + d11 v\u2217 y\u0307\u2217 + k11y \u2217 + k12\u03c8 + \u03b130y \u22173 + \u03b121y \u22172\u03c8 + \u00b7 \u00b7 \u00b7 + \u03b103\u03c8 3 + \u03b150y \u22175 + \u03b141y \u22174\u03c8 + \u03b132y \u22173\u03c82 + \u00b7 \u00b7 \u00b7 + \u03b105\u03c8 5 = 0, (2) \u03c8\u0308 + d22 v\u2217 \u03c8\u0307 + k21y \u2217 + k22\u03c8 + \u03b230y \u22173 + \u03b221y \u22172\u03c8 + \u00b7 \u00b7 \u00b7 + \u03b203\u03c8 3 + \u03b250y \u22175 + \u03b241y \u22174\u03c8 + \u03b232y \u22173\u03c82 + \u00b7 \u00b7 \u00b7 + \u03b205\u03c8 5 = 0, (3) where dots denote the derivative with respect to the dimensionless time t\u2217 and v\u2217 denotes the dimensionless running speed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001357_j.oceaneng.2019.106812-Figure18-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001357_j.oceaneng.2019.106812-Figure18-1.png",
        "caption": "Fig. 18. Comparison of cross section shape after rack and cylinder welding.",
        "texts": [
            " It can be seen that predicted welded zone is almost identical to the designed weld groove, and arc welding heat source a little bit penetrates the cylinder thickness. As mentioned above, this kind of temperature distribution will generate welding plastic strain in the outer surface and then influence the cross section shape of examined cylindrical leg structure. Appling the computed temperature profile as thermal loading, mechanical analysis during welding was also carried out to compute the plastic strains, and predict the welding deformation and residual stress. Concentrating on the cross section shape of cylinder, Fig. 18(a) shows that the left edge has a negative welding deformation in horizontal direction, which means that the left edge expands outward with respect to origin. In addition, Fig. 18(b) shows that the upper edge also has a negative welding deformation in upright direction, which means that the upper edge shrinks inward with respect to origin. Dealing with the symmetrical characteristic, Fig. 19 shows the cross section shape of examined cylindrical leg structure after rack and cylinder welding, in which original and deformed shapes were marked with yellow and orange colors. It is also can be seen that same deformation trends comparing with experiment observation. Although the computed welding deformation has identical tendency comparing with experimental observation, it is still necessary to compare the magnitude of computed welding deformation and measured data"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001712_j.ast.2020.105936-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001712_j.ast.2020.105936-Figure2-1.png",
        "caption": "Fig. 2. Illustration of the considered AHV.",
        "texts": [
            " Both RCS quantization and ACS nonlinearity are well handled. 2) In the back-stepping design for the altitude loop, no parameter update law is introduced, while the control structure is quite simple with low computational burden. In the rest of this paper, section 2 gives the considered AHV model, section 3 implements the control design for the velocity and altitude loops, section 4 presents the simulation study, and section 5 concludes this paper. The AHV under consideration is illustrated in Fig. 2, whose attitude is controlled by ACS and RCS, cooperatively. More specifically, the ACS suite contains a couple of elevators configured at the afterbody, while the RCS suite contains a set of small jet thrusters located near elevators. The function of RCS is to assist ACS especially when working at a high altitude. Note that RCS jet thrusters are not installed near the commonly suggested nose like space shuttles, because the forebody is too thin to install jet thrusters. Using the pulse-width modulation technique, RCS jet thrusters produce equivalent control forces, whose values are discrete with respect to the continuous RCS command"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003688_1.1304914-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003688_1.1304914-Figure2-1.png",
        "caption": "Fig. 2 The Euler angles and reference frames used to parameterize the precessional motion of the track",
        "texts": [
            "org/ on 04/28/20 We prescribe the rotational motion of the track such that a body-fixed axis normal to the track\u2019s plane precesses with a constant nutation angle u and a precessional rate c\u0307 . To describe the rotation, we introduce a right-handed orthonormal basis \u00ab of an inertial coordinate system comprised of the vectors E1 , E2 and E3 . Starting from a reference configuration where the normal to the track\u2019s plane is aligned with E3 , the rotation of the track may be specified by a 3-1-3 ~c, u, w! set of Euler angles, as shown in Fig. 2. Since we desire that the outer shell have a purely precessional motion that does not involve the body revolving about E3 , the final angle of rotation is w52c . In the resulting motion, a material point of the track returns to the same location after one full precession without revolving around E3 . We define a set of body-fixed vectors $et1 ,et2 ,et3% that form a right-handed orthonormal basis \u00abt which corotates with the track. Here, et1 and et2 lie in the plane that contains the track, while et3 is normal to the track\u2019s plane"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002218_j.mechmachtheory.2019.103771-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002218_j.mechmachtheory.2019.103771-Figure10-1.png",
        "caption": "Fig. 10. Transformation of different coordinate systems.",
        "texts": [
            " Then the lateral displacement function W ( \u0304x , t ) and the longitudinal displacement function V ( \u0304x , t ) of the beam element at a certain position x\u0304 , can be expressed as \u23a7 \u23aa \u23a8 \u23aa \u23a9 W ( \u0304x , t ) = \u2211 i f i ( \u0304x ) u i ( t ) ( i = 2 , 3 , 5 , 6 ) V ( \u0304x , t ) = \u2211 i f i ( \u0304x ) u i ( t ) ( i = 1 , 4 ) (3) where f i ( x ) is the type function, which is a function of the position coordinate x\u0304 . Another two coordinate systems are used to facilitate the modeling process: the fixed global coordinate system OXY and the rotary coordinate system Oxy whose axes are parallel to the corresponding axes of the local coordinate system A x y at each position of the mechanism. The relationships among the different coordinate systems are shown in Fig. 10 . In the moving process of the mechanical system, the coupled terms between the elastic deformations and the rigid-body movements can be ignored, since these coupled elastic deformations are usually very small. Then the following assumed motion relationships in the rotary coordinate system can be obtained as { \u02d9 u a = \u02d9 u r + \u02d9 u u\u0308 a = u\u0308 r + u\u0308 (4) where, \u02d9 u a , \u02d9 u r , \u02d9 u represent the absolute velocity, the rigid-body velocity and the velocity of the elastic deformation, respectively, and u\u0308 a , u\u0308 r , u\u0308 represent the absolute acceleration, the rigid-body acceleration and the acceleration of the elastic deformation, respectively",
            " Moreover, in the analysis process of single sub-mechanism in the following, the forces applied on this sub-mechanism generated by another two still sub-mechanisms are regarded as the external loads for accurate analysis. In this regard, the elastodynamic analysis by dividing whole mechanism into three sub-mechanisms is of reference values for practical design and control of the novel forging manipulator. (1) Elastodynamic modeling of lifting mechanism The lifting mechanism extracted from the main-motion mechanism of the forging manipulator is shown in Fig. 10 . The block and the piston rods of the lifting cylinder are regarded as seperate elements, the lifting arm AKCJ is divided into four elements, and the lifting mechanism is decomposed into six elements numbered by 1 \u00a9, 2 \u00a9, \u2026 6 \u00a9, whose nodes are marked by capital letters A, K, C, J, Q , and I . Meanwhile, subscript \u201cL\u201d is added to each parameter in the lifing mechanism to distinguish the parameters in different mechanisms. Corresponding to the generalized coordinates of the beam element in the local coordinate system, three generalized coordinates in the global coordinate system are set at each general node like A, J, C, K , respectively, which denote the elastic displacements in X and Y directions, as well as the elastic rotation angle at the beam node"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002552_jsyst.2020.3034993-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002552_jsyst.2020.3034993-Figure1-1.png",
        "caption": "Fig. 1. Configuration of the TRMS.",
        "texts": [
            " In order to prove capability of the proposed control algorithm, an experimental implementations are performed to the TRMS prototype. The adopted control approach permits to evade the modelling problems, to provide a best robustness and to get a desired trajectory tracking with better precision in attendance of wind effects. The rest of the article is arranged as follows, in Section II, the TRMS description with dynamic modelling are adopted. Section III presentsthe AFC design, experiment results are provided in Section IV. Finally, conclusions are given in Section V. II. TRMS DESCRIPTION As depicted in Fig. 1, the TRMS is an aero-dynamical system. Its comportment is like helicopter, composed of two perpendicular rotors actuated by dc motors, vertical rotor and horizontal rotor that are put on a beam which contains counter weight. The vertical rotor generates a hoist force permitting the beam to turn in main or vertical plane (pitch angle noted \u03c8), while the horizontal rotor permits the beam to turn in tail or horizontal plane (yaw angle noted \u03d5). The two rotors are commanded by varying the input voltage of the dc motors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000389_ijvd.2017.083418-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000389_ijvd.2017.083418-Figure3-1.png",
        "caption": "Figure 3 (a) Schematic picture of EBM; (b) wheel coordinate system according to ISO 8855, Sayers (1996) and c) definition of parameters at the EBM in movement",
        "texts": [
            " The brush tyre model is a physical model and provides the possibility of dealing with the force generation mechanisms of the tyre using mathematical representations to obtain solutions either analytically or, more elaborately, numerically through simulations. In the original brush tyre model (Hadekel, 1952), see Figure 2, and the models developed based on it (Pacejka, 2012; Sorniotti and Velardocchia, 2008) rolling resistance is either a predefined coefficient achieved from laboratory tests, or an empirical equation depending on other constants which are realised using experiments. However, in the extended brush tyre model (EBM) proposed here, see Figure 3(a), this coefficient is calculated by the model and can therefore be compared to experimental data. The EBM, as can be seen in Figure 3(a), similar to the basic brush tyre model where the rubber treads are assumed in form of individual bristles, uses finite number of bristles that touch the road plane. However, as shown, unlike the basic brush tyre model where the tyre width is disregarded, the EBM considers finite number of lines parallel to each other in order to resemble the tyre width and the bristles are spread over these lines along the contact patch. So the tyre is modelled as a rigid ring with bristles. In the following, the related kinematics of the tyre and the definitions which are used in the model are introduced",
            " The topology of the model is defined by the tyre radius, rim width, aspect ratio and side-wall inclination. In order to enable a realistic model of the tyreroad contact, the rubber properties have been implemented in all directions, i.e., longitudinal (x), lateral (y) and vertical (z) directions. Two sets of coordinate systems are used. The reference coordinate system is in accordance with the ISO definition, where its origin is attached to the contact patch, and the wheel local coordinate system is attached to the wheel centre, see Figure 3(b). The side-slip angle is defined as the deviation between the wheel heading and its longitudinal speed, where tan(\u03b1) = vy vx . (1) The energy losses in the side-wall are assumed to be negligible and thus not taken into account in this model. This is due to the fact that the volume of the side-wall is considerably less than the volume of the ply rubber and cord system of tyre. Thereby the loss modulus of the side-wall is much smaller than the ply rubber and tyre cords (Willett, 1973). Therefore, the carcass and belt construction of the tyre are assumed to have no damping and the belt is assumed as a rigid ring which the bristles are connected to",
            " As long as the bristles fulfil the friction ellipse condition, equation (8), the condition of the adhesion region is valid;( fx \u00b5s x fz )2 + ( fy \u00b5s y fz )2 \u2264 1. (8) This is made with the assumption that the wheel-spin around the vertical axis of the tyre is negligible. Under these conditions, the generation of adhesion force, fa, will be governed by the rubber deformation generated in correspondence with the slip velocity defined asvsx vsy  =  1 \u2212 sin(\u03d5) sin(\u03b3)  [ R \u03c9 ] \u2212 1 0 0 1 vx vy  (9) where R is considered to be the mean value of bristles\u2019 radii in the contact patch and \u03d5 determines the bristles\u2019 peripheral position, see Figure 3(c). As soon as the condition of friction ellipse is violated, the longitudinal or lateral slip occurs. In this situation, the bristles are entering into the sliding zone where the friction level and consequently the deflection change from the leading to the trailing edge. In this model neither the friction coefficient nor the contact force are derived based on the assumptions and methods used for adhesion region. Using the Coulomb friction it is not possible to express the slip phenomenon in the sliding region, since the sliding velocity and rubber forces will be influential in this case"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003720_s0039-9140(96)02040-1-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003720_s0039-9140(96)02040-1-Figure15-1.png",
        "caption": "Fig, 15. Microcell with working rotating disk electrode for batch injection analysis.",
        "texts": [
            " However, the system [99] is different from the others in its simplicity and compactness. In comparison with the microcell in Fig. 13, in the microcell in Fig. 14 steadier movement of the sample to the working electrode surface and a more intensive hydrodynamic effect have been achieved [99]. We have seen that the microcell with a static working electrode (Fig. 12) was used for batch injection in anodic stripping voltammetry and this design did not allow the use of a rotating working electrode for stripping voltammetric analysis. At the same time, the design in Fig. 15 [36, 37] with the rotating disk working electrode in the microcell for batch injection in anodic stripping voltammetry decreased the detection limit by approximately an order of magnitude. In Fig. 15, (1) is the inner microcompartment, (2) is the sample (200-300/tl), (3) is the body of the microcompartment, (4) is the outer large compartment, (5) is the supporting electrolyte, (6) is the channel o 0 6 0 cm < .= ~ 0 o ~T T @ @ N N g \"N o .~ o ~ ~ ~'~ \"\" ~ o . ~ ~ ~ = ' ~ ' ~ ~ ~ . ~ ~ ~ o ~ . ~ , ' , ~ 0 ~ \u00a2 ) N ~ ~ ~ ~~oo o . ~ ~ ~ . ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ oa.~.~ .~ ~ . ~ .a .~ .a .a ~ . ~ . ~ . ~ ~ ~ ~ = ~ ~ ~ ~ ' ~ ~ - . - . ~ . ~ . - ~ T @ o \"0 o o \"~ ~ = \".~ \u2022 - .~ ~ o 8 c~ ,-~ ~ ~ ~ 0 ~'~ ~ , "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001938_icra40945.2020.9196853-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001938_icra40945.2020.9196853-Figure1-1.png",
        "caption": "Fig. 1. Illustration of the structure and dynamic model of the snake robot. The top figure shows link i, its axes, torques, mass, and length. The bottom figure shows an example of a 5-link snake.",
        "texts": [
            " The dynamics of snake locomotion are well studied in the literature, as a dynamic model is often used to validate or tune a proposed gait. In our work, the dynamic model is also used to generate the trajectory, via optimization with respect to a given cost function. Before describing the gait synthesis method, we will thus describe the dynamic model we use, which follows a standard time-stepping method based on a Newton-Euler formulation with explicit integration for a 2D n-link kinematic chain. We employ the model depicted in Fig. 1, where q0 is the angle of the first link with respect to the world, and qi denotes the relative angle of link i with respect to the previous link (counterclockwise); \u03c4i is the motor torque applied to joint i; and p0 = [x0,y0] is the head position in the global coordinates. The length of link i is li and its mass is denoted by mi. We assume uniform distribution of mass so the center of gravity is located in the middle of each link; gravity is pointing into the page. Each link has a local coordinate frame based at the proximal end, with the yi axis pointing in the longitudinal direction and the xi axis in the transverse direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000319_icrom48714.2019.9071797-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000319_icrom48714.2019.9071797-Figure1-1.png",
        "caption": "Fig. 1. Schematic of a 3-DOF suspended CDPR. All of the anchor points are in a horizontal plane.",
        "texts": [
            " However, this will not be possible in the case of underactuated robots. In other words, the main reason we incorporate interconnection matrix is that it is possible to apply the proposed method for stabilization of underactuated robots with input constraint. IDA-PBC method alongside ensuring input constraint is our major future purpose. In this section we apply the preceding design methodology to a 3-DOF cable driven robot with three actuators through the simulation. Schematic of this robot is shown in Fig. 1. Configuration variables of the robot are X = [x, y, z]T . Dynamic parameters are considered as follows M = mI3 V = mgz. where m = 4Kg is the mass of end-effector. Jacobian matrix of this robot is [5]: Ja = x\u2212xa1 l1 y\u2212ya1 l1 z\u2212za1 l1 x\u2212xa2 l2 y\u2212ya2 l2 z\u2212za2 l2 x\u2212xa3 l3 y\u2212ya3 l3 z\u2212za3 l3  , in which lis are the length of cables computed as follows l1 = \u221a (x\u2212 xa1)2 + (y \u2212 ya1)2 + (z \u2212 za1)2 l2 = \u221a (x\u2212 xa2)2 + (y \u2212 ya2)2 + (z \u2212 za2)2 l3 = \u221a (x\u2212 xa3)2 + (y \u2212 ya3)2 + (z \u2212 za3)2, and kinematic parameters are xa1 = xa2 = \u2212xa3 = b/2 = 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure41.2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure41.2-1.png",
        "caption": "Fig. 41.2 Base station design",
        "texts": [
            " Preliminary testing includes battery charge-level check, power plant test run, inertial navigation sensor system health check and, finally, ensuring the presence of the specified number of satellites to flight along the route. The developed diagram enables composition of scenarios of interactions between UAV and base station. These scenarios can be applied in algorithm design for UAV and base station control. Based on the presented list of functional features, a 3D model of base station was designed. This model provides for visualization of possible scenarios andd interactions between the UAV and the base station. This model is depicted in Fig. 41.2 and consists of two levels. The base station housing is fabricated of aluminum profiles and PVC panels. These materials were chosen because of the requirement to minimize the overall structure weight. The top-level elements of the station are mounted above the housing; there are essentially the roof and meteo-sensors, which can read data on environmental conditions. Thermometers, barometers, anemometers and hygrometers can be used as such sensors. The bottom level of the station is established inside the housing and includes a landing site, landing site elevation unit, battery replacement unit, station control unit, station power supply unit, as well power distribution system",
            "3 illustrates how the centeringmechanism shifts along the rails laterally. When themechanism touches theUAVbase, theUAV shifts laterally (see Fig. 41.3b). Figure 41.3c shows the final step of mechanism and UAV motion toward the center of the site. Battery replacement is fulfilled without aid of technical vision or other UAV discovery methods. After the UAV has been centered on the landing site, the battery will be extracted, and battery replacement is performed, afterward the battery replacement mechanism shifts to point 3 (see Fig. 41.2). The final step of battery replacement (see point 4 in Fig. 41.2) is battery insertion into dock station. Installation of the recharged battery into the UAV is performed in reverse order (see points 1\u20134 in Fig. 41.2). After completion of functional block \u00abUAV battery replacement\u00bb, the station switches to mode \u00abBattery storage\u00bb until the next replacement mission. Two-way data transfer module enables data exchange at distance up to 50 kmwith operational frequency grids of 2.3/2.4/2.5 and 5.2\u20135.8 GHz. Maximum data transfer speed is 12Mbit/s. Video stream can be transferred with resolution of 1920\u00d7 1080p, frequency 30 frames/sec. and delay 125 ms. To obtain publicly available meteodata, a 4G modem is used. To implement functional blocks \u00abUAV landing\u00bb, \u00abUAV storage\u00bb and \u00abUAV takeoff\u00bb, a retractable roof was developed (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002675_0142331220983637-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002675_0142331220983637-Figure1-1.png",
        "caption": "Figure 1. Structure of the two-link manipulator.",
        "texts": [
            " 0, 0, x= 0, 1, x\\0: 8< : According to this definition, the following equation holds dsign(x) dx = njxjn 1: 8x 2 Rn and n 2 R\u00fenf0g, the vector field function x 7!sign(x) is defined as sign(x)= sign(x1), sign(x2), . . . , sign(xn)\u00bd T : For any selected vector r= frign i= 1, where ri 2 R\u00fenf0g, define the matrix Lr(l)= diagflrign i= 1,8l 2 R\u00fenf0g: Let I 2 Rn 3 n denote the identity matrix, and O 2 Rn 3 n denotes the null matrix. For the matrix A= \u00bdaij n 3 n, denote jAj= det(A) and the matrix norm is given by k A k = max x 6\u00bc0 Axkk xkk : The two-link manipulator (Zheng and Chen, 2018) with the structure shown in Figure 1 is considered in this paper. According to the analysis in Zheng and Chen (2018), the dynamic model of the two-link manipulator is M(u)\u20acu+C(u, _u) _u+G(u)= t + td , \u00f01\u00de where u= u1, u2\u00bd T denotes the joint positions, _u and \u20acu represent the velocity and acceleration, respectively, M(u) 2 R2 3 2 is the symmetric inertial matrix, C(u, _u) 2 R2 refers to the centrifugal-Coriolis matrix, and G(u) 2 R2 denotes the influence of gravity. The corresponding items are represented as follows (Zheng and Chen, 2018) M(u)= p1 + p3 + 2p2 cos u2 p3 + p2 cos u2 p3 + p2 cos u2 p3 , C(u, _u)= p2 _u1 sin u2 2p2 _u1 sin u2 p2 _u1 sin u2 0 , G(u)= p4 sin u2 + p5 sin (u1 + u2) p5 sin (u1 + u2) , in which p1 =(m1 +m2)l 2 1, p2 =m2l1l2, p3 =m2l2 2, p4 =(m1 +m2)gl1, p5 =m2gl2, where m1, m2 are the masses, l1, l2 are the lengths of the two links, t = t1, t2\u00bd T denotes the input torque, and td = \u00bdtd1, td2 T is unknown external time-varying disturbance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000215_j.neucom.2019.10.088-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000215_j.neucom.2019.10.088-Figure1-1.png",
        "caption": "Fig. 1. Underactuated mechanical system.",
        "texts": [
            " Remark 6. By combining the technique of adding a power integrator, graph theory and Young\u2019s inequality together, we propose a new distributed controller design approach for system (1) . From the design procedure developed in Section 3.1 , it can be observed that our control scheme can effectively deal with the unstabilizable linearizations in each agent and the nonlinear interactions among agents simultaneously. 4. Applications to mechanical MASs Consider the underactuated mechanical systems [24] (see Fig. 1 ) described by \u03b8\u0308 = g l sin \u03b8 + k s m 2 l (x \u2212 l sin \u03b8 ) 3 cos \u03b8, x\u0308 = \u2212 k m 1 x \u2212 k s m 1 (x \u2212 l sin \u03b8 ) 3 + v m 1 . (38) The definitions of \u03b8 , l, k s , m 1 , m 2 and v can be found in [24] . Introduce the following coordinate transformations x 1 = \u03b8, x 2 = \u02d9 x1 , x 3 = (x \u2212 l sin \u03b8 ) 3 \u221a k s m 2 l cos \u03b8, x 4 = \u02d9 x3 Please cite this article as: W. Li, L. Liu and G. Feng, Containment cont unstabilizable linearizations, Neurocomputing, https://doi.org/10.1016/j. n (x 1 , x 2 , x 3 , x 4 ) \u2208 (\u2212(\u03c0/ 2) , \u03c0/ 2)) \u00d7 R 3 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002572_j.apm.2020.12.020-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002572_j.apm.2020.12.020-Figure2-1.png",
        "caption": "Fig. 2. Deformation description of an arbitrary flexible body during the movement.",
        "texts": [
            " The lightweight tracking mechanism is prone to undamped continuous elastic vibration in the space tracking motion, which will significantly reduce the dynamic tracking accuracy of the mechanism. Therefore, the flexible multi-body dynamic modeling and dynamic characteristics analysis of the tracking mechanism in this paper have some theoretical significance and lay the foundation for the subsequent vibration control system design. Based on the floating coordinate frame, that is, the elastic deformation of flexible body is described in its body-fixed coordinate system. Without losing generality, Fig. 2 shows the deformation description of any arbitrary flexible body j in the process of large-scale rigid body movement. Shear stress of the beam is not the focus and ignored in this paper. The Timoshenko beam element which considers the effect of shear stress is mostly used for elasto-dynamic modeling of short and thick beams. However, limbs of the parallel tracking mechanism in this paper are slender bars with their slenderness ratios bigger than 10. The effect of shear stress has little influence in large-scale flexible body motions and can be ignored [23] ",
            " 4 , the node generalized coordinates array of any element i ( i = 1,2, , n j ) in the flexible body j is u f i, j = ( u j 1 ,i u j 2 ,i u j 3 ,i u j 4 ,i u j 5 ,i u j 6 ,i u j 7 ,i u j 8 ,i u j 1 ,i +1 u j 2 ,i +1 u j 3 ,i +1 u j 4 ,i +1 u j 5 ,i +1 u j 6 ,i +1 u j 7 ,i +1 u j 8 ,i +1 )T (3) where u j 1 ,i ( u j 1 ,i + 1 ), u j 2 ,i ( u j 2 ,i + 1 ), u j 3 ,i ( u j 3 ,i + 1 ) denote the elastic displacements of the left (right) node for element i along x\u0304 i , y\u0304 i , z\u0304 i axis, respectively; u j 4 ,i ( u j 4 ,i + 1 ), u j 5 ,i ( u j 5 ,i + 1 ), u j 6 ,i ( u j 6 ,i + 1 ) denote the elastic angles of the left (right) node for element i around x\u0304 i , y\u0304 i , z\u0304 i axis, respectively. u j 7 ,i ( u j 7 ,i + 1 ) and u j 8 ,i ( u j 8 ,i + 1 ) are elastic curvatures around y\u0304 i axis and z\u0304 i axis of left (right) node of element i. Three dimensional isoparametric beam element with curvature is used for dynamic modeling for its high accuracy. As shown in Fig. 2 , the deformation displacement vector \u03b4P i, j of an arbitrary point P i, j of element i in the flexible body j under frame O j \u2212 x\u0304 i \u0304y i \u0304z i can be expressed as \u03b4P i, j = \u23a1 \u23a3 u j i ( \u0304x i , y\u0304 i , \u0304z i , t ) v j i ( \u0304x i , y\u0304 i , \u0304z i , t ) w j i ( \u0304x i , y\u0304 i , \u0304z i , t ) \u23a4 \u23a6 = N i, j ( \u0304x i , y\u0304 i , \u0304z i ) u f i, j (4) where N i, j ( \u0304x i , \u0304y i , \u0304z i ) represents the shape function matrix of three-dimensional isoparametric beam element i in the flexible body j , and N i, j ( \u0304x i ) = [ N i, j, 1 ( \u0304x i , y\u0304 i , \u0304z i ) N i, j, 2 ( \u0304x i , y\u0304 i , \u0304z i ) N i, j, 3 ( \u0304x i , y\u0304 i , \u0304z i ) ] = \u23a1 \u23a3 f j i, 1 f j i, 2 f j i, 3 0 f j i, 4 f j i, 5 0 0 0 f j i, 6 0 f j i, 7 0 f j i, 8 0 f j i, 9 0 0 f j i, 10 f j i, 11 f j i, 12 0 f j i, 13 0 f j i +1 , 1 f j i +1 , 2 f j i +1 , 3 0 f j i +1 , 4 f j i +1 , 5 0 0 0 f j i +1 , 6 0 f j i +1 , 7 0 f j i +1 , 8 0 f j i +1 , 9 0 0 f j i +1 , 10 f j i +1 , 11 f j i +1 , 12 0 f j i +1 , 13 0 \u23a4 \u23a6 (5) where x\u0304 i , y\u0304 i , z\u0304 i are the coordinates of point P i, j in the frame O j \u2212 x\u0304 i \u0304y i \u0304z i "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001900_s0263574720000806-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001900_s0263574720000806-Figure4-1.png",
        "caption": "Fig. 4. Model of the future: (a) space\u00d7 time conservative model (partially represented for visualization purposes) for a scenario with three fixed objects and one moving object. Every point of the limit of the field-of-view and occlusions (unseen objects) is modelled as a disc that grows as time passes (i.e., a cone in space\u00d7 time). The fixed object remains constant over time (i.e., a vertical line in space\u00d7 time). The moving objects are modelled according to their future behaviour (i.e., a curve in space\u00d7 time) if it is available and reliable, otherwise it is treated as an unseen object and modelled as a growing disc. (b) How FOV shrinks as time passes (for a time t1 greater than the sensing time), with \u03b4FOV the limit of the field-of-view and FOV the subset of WS perceived by A.",
        "texts": [
            " Given Property 1, Property 2 guarantees that at each planning cycle, a p-safe partial trajectory is available. Another important element regarding safety issue is the model of the future. To consider all possible future motions of a given moving object with an unknown future behaviour (seen or unseen), the model used herein is conservative: given an upper bound of the velocity of objects, every point of the limit and outside the field-of-view is modelled as a disc that grows as time passes, that is, a cone in space\u00d7 time (see Fig. 4).2 Regarding this model, it is necessary to answer the question how far into https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574720000806 Downloaded from https://www.cambridge.org/core. Cornell University Library, on 12 Sep 2020 at 17:55:55, subject to the Cambridge Core terms of use, available at the future reasoning is done? This model must still valid over a limited interval of time to guarantee safety, hence, the following property: Property 3 (Future model validity). The model of the future has to remain valid over the time interval [tk, tk+2 + Th], for every cycle k, with tk+2 = tk + 2 \u2217 \u03b4cycle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002755_1077546321998559-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002755_1077546321998559-Figure2-1.png",
        "caption": "Figure 2. Deep groove ball bearing model.",
        "texts": [
            " By comparison with the results of the static model, the significant influence of gear meshing on the internal load and bearing fatigue life is demonstrated. It is well known that an external load applied to the bearing is distributed among the rolling elements. However, few researchers had considered the bearing internal load in the gear dynamic model. Figure 1 displays the dynamic model of a gear pair, with the bearing model introduced as a rigid body or a linear spring. As a result, the internal load distribution cannot be acquired using a dynamic model. As shown in Figure 2, the dynamic model of the deep groove ball bearing is established to investigate the bearing load distribution in actual motion. The bearing\u2019s outer ring is fixed constraint, and the inner ring rotates with the center shaft. The instantaneous orbital position angle of the j-th rolling element can be written as (Liew and Lim, 2005) \u03c8j\u00f0t\u00de\u00bc V 2 1 D dm t\u00fe2\u03c0 Z \u00f0 j 1\u00de, \u00f0 j\u00bc 1;2,\u2026,Z\u00de (1) whereV is the mean shaft rotational speed, D is the element diameter, and dm is the bearing pitch diameter. Z is the number of elements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001312_tie.2019.2952780-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001312_tie.2019.2952780-Figure1-1.png",
        "caption": "Fig. 1. A conventional hydraulic powered lathe machine along with a schematic diagram of the headstock showing the internal components [5].",
        "texts": [
            " (Corresponding Author: J. Chang, e-mail: cjhwan@dau.ac.kr). D. Jang, H. Shin, S. Paul and J. Chang are with the Mechatronics System Research Lab, Electrical Engineering Department, Dong-A University, Busan, South Korea, 49315. Yongseon Yun is with the KHAN Work Holding Co.,Ltd., Changwon, South Korea. ease of operation, accuracy and repeatability. Beacuse of their popularity a good amount of works have been published recently which address the structrual design, machining and control of the CNC lathe [6]-[7]. Fig. 1 shows a conventional CNC lathe with schematic diagrams of the components of the headstock of the machine. In the CNC lathe of Fig. 1, two major operations are performed to machine a workpiece, namely, the driving of the clamping system and the rotation of the spindle to rotate the clamping system holding the workpiece. Conventionally, the clamping force for grasping the workpiece is provided by the hydraulic pressure generated using a hydraulic cylinder. Whereas the rotary motion of the spindle is achieved using the electric motor attached to the headstock of the CNC lathe which transmit the rotary motion through the pulley-belt mechanism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000809_ecc.2016.7810345-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000809_ecc.2016.7810345-Figure1-1.png",
        "caption": "Fig. 1. Several trajectories of the system",
        "texts": [
            " In fact, if there are no discontinuities in the vector field, the conditions in Theorem 2 are reduced to those in [12]. In particular, the inequality conditions (17) on the discontinuity surfaces vanish if f(x) and \u03c1(x) happen to be continuous there. To illustrate the main result, here we show a simple example of a Lyapunov density of Theorem 2. Let x \u2208 R 2 and consider the system (1) with f1(x) = [ 1 \u22124 4 \u22122 ] x, f2(x) = [ \u22123 \u22124 12 2 ] x, where X1 = {x \u2208 R 2 : x2 \u2265 0}, X2 = {x \u2208 R 2 : x2 \u2264 0} [3], [8]. Figure 1 shows trajectories of this discontinuous system for several initial values. This system meets the assumptions stated in Section II. We can verify that the Lyapunov density given by \u03c1i(x) = 1/(x\u22a4Pix) 6, i = 1, 2 with positive definite matrices P1 = [ 1 \u22120.332 \u22120.332 0.853 ] , P2 = [ 1 0.224 0.224 0.333 ] satisfies the conditions of Theorem 2. In fact, each \u03c1i(x) is obviously positive if x 6= 0 and so is [\u2207 \u00b7 (fi\u03c1i)](x), as shown in Figure 2. In this example, we have X12 = X21 = {x \u2208 R 2 : x2 = 0}, b12(\u03be) = b21(\u03be) = [ \u03be 0 ]\u22a4, h12 = [ 0 1 ]\u22a4"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002046_icem49940.2020.9270988-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002046_icem49940.2020.9270988-Figure3-1.png",
        "caption": "Fig. 3 3D FEA structural model of the stator with tooth tips loads in yellow and outer yoke response nodes in green",
        "texts": [
            " It consists in computing separately electromagnetic operational loads and structural response under unit-magnitude loads (Frequency Response Functions, FRF) and then making the synthesis based on structural model linearity assumptions to get operational vibrations. This process is faster compared to a direct computation approach and gives more physical insights. 2) Mechanical model FRF are obtained from a 3D FEA model using MANATEE automated coupling with Hypermesh/Optistruct [10]. To compute FRF, unit-magnitude force waves are applied on stator tooth tips and displacement are taken for all nodes belonging to stator yoke outer surface, as illustrated in Fig. 3. The stator yoke is clamped at lower end, the upper one between free. Stator material properties are given in Table 1. Reduced damping is set to 2%. Stator windings are not included in the model to simplify the structural model. To perform EVS with uneven magnetization assumption, FRF are computed for any wavenumbers between -8 and 8 and in both radial and tangential directions. Wavenumbers for which absolute value is larger than 8 are neglected in the following. They should not radiate significant airborne noise due to high yoke stiffness",
            " TABLE 3 ORDERS AND NATURAL FREQUENCIES OF STRUCTURAL MODES Mode identification (m,n) (m: circumferential order; n: longitudinal order) Natural frequency (Hz) Yoke \u201cbreathing\u201d mode (0,0) 5450 Yoke \u2018bending modes (1,0) 963 & 970 Yoke \u201covalization\u201d modes (2,0) 995 & 1000 Radial modes (3,0) 1732 & 1737 Radial modes (4,0) 2698 & 2703 Tooth rocking modes (8,0) 4513 & 4519 2) Displacement FRF, r = 0,2,8 RMS value of displacement FRF under radial (respectively tangential) unit-magnitude force waves of wavenumber r = 0,2,8 are illustrated in Fig. 10 respectively Fig. 11). RMS displacement value is obtained by averaging the squared displacement value for each node of the outer yoke surface (cf. Fig. 3). In even magnetization case, vibrations are only due to FRF associated to wavenumbers r = -8,0,8. Therefore, main resonance peaks are due to radial pulsating (r = 0) force waves resonating with breathing mode around 5450 Hz and tangential rotating forces (r = \u00b18) resonating with tooth rocking mode around 4500 Hz. Fig. 10 FRF for radial unit-magnitude force waves FR F di sp la ce m en t [ m /(N /m 2 )] Tooth rocking mode contribution is overestimated in the mechanical response since stator winding stiffness is not modelled"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002533_j.tws.2020.107247-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002533_j.tws.2020.107247-Figure11-1.png",
        "caption": "Fig. 11. Definition of local coordinate system of hollow truss.",
        "texts": [
            " According to the failure modes map, by increasing the temperature or increasing d/l, the hollow truss lattice core is sufficient to be damaged by buckling. When the length of hollow truss is fixed, one method of avoiding the hollow truss buckling is to decrease t/d or increase t/d at \u201ctrough\u201d. To further explore the failure mechanism, FEM is employed to obtain the stress distribution of hollow truss at different temperature. Firstly, the local coordinate system is established for hollow truss as shown in Fig. 11. In this local coordinate system, a specific path ( op\u0305\u2192) along the length direction is defined. s = \u03b6/l denotes the relative location of an arbitrary point along hollow truss, and \u03b8 represents the angle of rotation around z axis of hollow truss wall. The mises stress distribution is obtained from FEM as shown in Fig. 12. The nodal stresses located in the defined path are selected and plotted in the curves as shown in Fig. 12(a). Clearly, the stress curves of hollow truss with t/d of 0.02 exhibit fluctuates along the length direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of internal gear slicing",
        "texts": [
            " The working relief angle is then calculated and analyzed for the purpose of avoiding interference. The faces of the designed cutter, especially the rake face, are complex and curved, leading to difficulty in manufacturing. In view of this, a point-grinding method is proposed for the curved faces to make the cutter design into a reality. At last, a gear slicing experiment is carried out to verify the validity of the proposed cutter design and manufacturing methods. When slicing the workpieces with internal teeth, the cutter is located inside the workpiece as shown in Fig. 1. There is an intersection angle \u03b3 between the cutter shaft and the workpiece shaft. The direction of the cutter tooth is required to be consistent with that of the workpiece slot. The cutter rotates with the workpiece at a certain speed ratio, and the workpiece feeds axially at the rate f. As the basis of mathematical derivation, the cutter coordinate system S2 (o2, x2, y2, z2) is first established as shown in Fig. 2, where axis z2 coincides with the centerline of the cutter. On the premise that the cutting edge is error-free, the rake face can be constructed in two ways: by a plane that is easy to manufacture or by a curved surface that can ensure consistent and reasonable working rake angle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001175_s0005117919090121-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001175_s0005117919090121-Figure2-1.png",
        "caption": "Fig. 2. A schematic image of a quadrotor.",
        "texts": [
            " (1), let\u2019s introduce two equations: I\u03c9\u0307 = \u239b \u239c\u239c\u239d L(F2 \u2212 F4) L(F3 \u2212 F1) \u03be(F1 + F3 \u2212 F4 \u2212 F2) \u239e \u239f\u239f\u23a0\u2212 \u03c9 \u00d7 I\u03c9, (2) H\u0308 = 1 m ( 1\u2212 2q2x \u2212 2q2y ) (F1 + F2 + F3 + F4)\u2212 g. (3) The Eq. (2) is the Newton\u2013Euler equation for a rotating body (see [3]). Here Fi is the thrust force (N) created by the ith propeller, I is the matrix of inertia moments (kg \u00d7m2), assumed to be diagonal, L is the distance from the center of mass to the propellers (m), and \u03be is the ratio of the thrust force of the propeller with the reactive moment (m). The correspondence between the propeller numbers and their positions and directions of rotation can be seen in Fig. 2. Equation (3) describes the quadrotor dynamics along the vertical axis Z directed against gravity (movement in the horizontal planeXY is omitted for simplicity, since in this paper we do not observe or attempt to control the position of the robot in this plane). Here H is the coordinate of the robot along the Z axis (m), m is the mass of the robot (kg), g is the gravitational acceleration (m/s2). The multiplier ( 1\u2212 2q2x \u2212 2q2y ) is equal to the cosine of the angle between the Z and Z \u2032 axes and determines how much the total thrust of the quadrotor affects its acceleration in height"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000225_ecce.2019.8912859-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000225_ecce.2019.8912859-Figure1-1.png",
        "caption": "Fig. 1 Examples of non-adjacent broken bars: end ring of (a) 6.6 kV, 500 kW, 10 pole coal crusher and (b) 380 V, 190 kW, 4 pole fuel pump induction motors with location of broken bars shown.",
        "texts": [
            " It has been shown that testing the motor when the rotor slip is high (such as standstill or starting transient) is immune to most of the false indications [6], [10]-[13]. It is common for bars adjacent to the damaged bars to break because current is redistributed to neighboring bars increasing the operating stresses [1]-[2]. However, cases of non-adjacent broken bars are also observed for applications with frequent transient stresses [6]. Examples of non-adjacent broken bars that have been observed by the authors are shown in Fig. 1 for 6.6 kV, 500 kW, 10 pole coal crusher, and 380 V, 190 kW, 4 pole fuel pump induction motors. The end rings cut off from the rotor cage clearly shows that bar breakages can be randomly spread out. MCSA based on monitoring the 2x slip frequency sidebands of the fundamental frequency, fs, = (1 \u00b1 2 ) \u2219 , (1) is very effective for detecting adjacent broken bars (k is an integer, and s is the rotor slip). However, it has been shown in [7], [14]-[20] that rotor faults are difficult to detect if the location of non-adjacent broken bars is such that the electrical asymmetry of the rotor observed from the stator winding is canceled. For example, the electrical asymmetry seen from the stator is canceled if 2 broken bars are separated by 90o (electrical) or half pole pitch (or if 3 broken bars are separated by 120o (electrical), etc\u2026). Therefore, it is likely for the rotor fault indicator to be significantly reduced and not proportional to the number of broken bars, if the bars are spread out as in the cases of Fig. 1. This causes MCSA based on fbrb to produce a false negative indication where the fault cannot be observed, and potentially lead to a forced outage without prior warning. Commercially available electrical tests including MCSA, single phase rotation test, and starting current analysis rely on indirectly observing the electrical asymmetry produced by broken bars from the stator winding. Therefore, they cannot 978-1-7281-0395-2/19/$31.00 \u00a92019 IEEE 7019 detect non-adjacent broken bars that interact to cancel the electrical asymmetry [6]-[7], [14]-[20]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002741_tim.2021.3055291-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002741_tim.2021.3055291-Figure5-1.png",
        "caption": "Fig. 5. (a) Experimental setup. (b) Outer ring faulty bearing. (c) Inner ring faulty bearing.",
        "texts": [
            " Step 5: Based on the constructing superimposed ITFR and the parameters of rolling bearing, we can achieve the target of fault diagnosis for rolling bearing under nonstationary conditions, and the fault location can be determined. In this section, the vibration signals of rolling bearing with different health conditions under nonstationary speeds are used to verify the proposed method. Moreover, some state-of-the-art TFA methods are used for comparison. The data that we used for verification are a public data set, which was provided by the University of Ottawa [36]. The experiment was completed on the SpectraQuest machinery fault simulator (MFS-PK5M), and the experimental setup is shown in Fig. 5(a). The motor drives the shaft, and the ac drive controls the rotational speed. A healthy bearing (left, ER16K ball bearing) and an experimental bearing (right) are used to support the shaft. The outer ring faulty bearing and inner ring faulty bearing are shown in Fig. 5(b) and (c), respectively. In order to collect vibration data, an accelerometer (ICP accelerometer, Model 623C01) was placed on the housing of the experimental bearing. Moreover, an incremental encoder (EPC model 775) was used to measure the shaft rotational speed. The final data are collected by the NI data acquisition boards (NI USB-6212 BNC). In addition, the parameters of ER16K ball bearing are given in Table II. According to these parameters, the characteristic frequencies of faults in different parts of the bearing can be calculated"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002220_j.triboint.2020.106167-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002220_j.triboint.2020.106167-Figure5-1.png",
        "caption": "Fig. 5. Geometric model of Experiment 4 with a 100-N load measured using the CT system (grey zones correspond to the contact zone).",
        "texts": [
            " After applying the squeeze load (5, 50, 100, 200, 350, and 500 N for contact pairs 1\u20134; 5 and 125 N for contact pairs 5 and 6) by screwing the load bolt, the apparatus was set up on the objective table of the CT equipment. Then, the X-rays emitted by the radioactive source were passed through the sample and detected and recorded by the detector. The objective table was then rotated to a new position by a preset number of degrees, and the sample was irradiated by additional X-rays. The detection procedure continued until the objective table had been rotated by 360\ufffd. The data recorded in each position were processed based on tomographic reconstruction and other methods [32] to build a geometric model (Fig. 5). Specifically, this model was output in stereolithography (STL) format and contained triangulated surfaces that defined the external surfaces of the contact solids, as well as point clouds defining the corners of the triangulated surfaces. Stereolithography models are computer-aided design models that are widely used to describe the external surfaces of objects. In a contact pair, gaps exist between rough surfaces in the non-contact region, while the two solids come into contact in the contact region via the squeeze pressure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002746_j.triboint.2021.106920-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002746_j.triboint.2021.106920-Figure8-1.png",
        "caption": "Fig. 8. Circular grooved texture.",
        "texts": [
            " Therefore, the verification of the irrelevance between the time steps number and calculation results need to be carried out and the specific method is increasing the time steps number, until the time steps number increased, the calculation result changed very little. Through this process, it can be found that the calculation time of each step should be less than or equal to 0.00005s if the accuracy of calculation is to be guaranteed. Therefore, the time steps will number vary depending on the case, when the groove spacing is 0.04 mm, 0.06 mm and 0.08 mm, the time steps number, n, are 8, 12 and 16, respectively. Fig. 9 shows the simulation results of the reciprocating rod seal with a circular grooved texture rod (see Fig. 8). Because of the grooves, a C. Xiang et al. Tribology International 158 (2021) 106920 small crest and trough of fluid pressure occurs at the corresponding position of each groove, and the asperities contact pressure is low where a groove is present and high where no groove exists. The working process of the sealing system with a circular grooved texture rod is a cyclic transient process. Therefore, when the rod moves to different positions, the pressure results differ and the peak point changes with relative position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002765_s11071-021-06327-0-Figure20-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002765_s11071-021-06327-0-Figure20-1.png",
        "caption": "Fig. 20 Dynamical model of a tracked vehicle",
        "texts": [
            " At each moment, the brightest position on the y-axis corresponds to the main frequency. The red curve represents the excitation frequency caused by the polygon effect [8] calculated from the following equation. ft \u00bc v=PT \u00f035\u00de This excitation is mainly generated by the impact of the sprocket on the moving tracks [9]. It is seen that the two are in good agreement. Therefore, the track is mainly excited by a high-frequency caused by the polygon effect during the running of tracked vehicles. Based on MSTMM, a nonlinear dynamical model of a tracked vehicle is developed, as shown in Fig. 20. The proposed contact algorithm is adopted in the dynamical simulation. Its step size is 10\u20133 s. It took about 25 min on a computer with core i5 dual-core 1.8 GHz processors and 8 GB of RAM. The vehicle is accelerated from 0 to 7.4 m/s within 10 s. Then it is driven at this constant speed. The corresponding excitation frequency of the polygon effect is 54 Hz. Figure 21 illustrates a simulation result of the power spectral density (PSD) for the corresponding vibration acceleration. Comparing frequency domains between the simulation result of the acceleration signal and the test data, their main peak frequencies are consistent"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001526_s00170-020-04987-7-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001526_s00170-020-04987-7-Figure6-1.png",
        "caption": "Fig. 6 Calibration and sensitivity study artifacts. a Printed plate. b Example of the CMM-measured distortions (mm)",
        "texts": [
            " For simulations, the following properties were either directly taken or calculated from the material data sheet provided by EOS [30] (EOS AlSi10Mg powder): the as-built in-plane Young\u2019s modulus, yield strength, ultimate strength and elongation at break, and strain hardening factor (Table 2). Poisson\u2019s ratio was taken from the MatWeb site [31]. After the parts were printed, a thin layer of talcum powder was used to reduce part surface reflection. In doing so, the potential point cloud density was increased to ensure the best measurement. CMM scans of the entire build plate (with no postprocessing treatment of the parts) were then carried out. In- plane distortions were measured on each of the calibration artifacts. Figure 6 a presents the calibration and the sensitivity study artifacts lying on the build plate, and Fig. 6 b presents an example of CMM-measured distortions. Point clouds of all the printed parts were obtained using aMetris LC50 laser scan mounted on a Mitutoyo CMM (accuracy \u2264\u00b1 7.5 \u03bcm at 95% confidence level). Distortion measurements were carried out using a metrological software certified by the National Metrology Institute of Germany (PTB) PolyWorks\u00ae v.11 (InnovMetric, Quebec, Canada), and point cloud post-processing, using statistical analysis software Matlab\u00ae 2017b (MathWorks, Natick, MA, USA) and Minitab\u00ae v"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001607_0954407020909663-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001607_0954407020909663-Figure1-1.png",
        "caption": "Figure 1. Entire vehicle model.",
        "texts": [
            " By improving the understeer of the vehicle, the optimization of minimum time handling and stability was realized. Finally, the simulation analysis was carried out to compare the minimum timehandling and stability performance before and after optimization to verify the effectiveness of the optimization. Minimum time-handling and stability analysis of the vehicle The seven subsystems of a vehicle, including suspension, car body, and tire, are constructed using the ADAMS software. After proper simplification, Figure 1 shows that the subsystems are connected based on the communicator to constitute an assembly model of the finished vehicle. The ADAMS model has been validated in Lixia et al.14 Simulation analysis of double-lane change With path as the objective, optimal control is realized by compiling driver files based on the optimal path, which is obtained through the inverse dynamics study method of minimum time handling under the ADAMS simulation setting.14 Corresponding geometric stake and geometric constraints of the vehicle body, as well as the corresponding simulation environment of doublelane change, are created in ADAMS according to the GB6323-94 technical requirements"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001357_j.oceaneng.2019.106812-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001357_j.oceaneng.2019.106812-Figure10-1.png",
        "caption": "Fig. 10. Design of rigidity reinforcement solution.",
        "texts": [
            " In detail, cross section of examined cylindrical leg structure was transformed from circle to elliptic as: the distance between upper edge and lower edge in upright direction shrink about 20 mm, and the distance between left edge and right edge in horizontal direction expand about 7 mm this kind of geometrical change caused by welding will influence the interaction between cylindrical leg structure and working platform, even the operation performance and efficiency for whole jack-up rig. In detail, Fig. 9 shows the cylinder straightness of four positions (right, left, down and up points as indicated in Fig. 6) at each measured section after welding, while the negative deformation means inward shrinkage and positive deformation means outward expansion. Currently, rigidity of cylindrical leg structure was improved with inner rib ring as shown in Fig. 10 (a) or pillar stiffener as shown in Fig. 10 (b) to avoid the welding distortion and ensure the fabrication accuracy. Although this practical approach can prevent the welding distortion as shown in Fig. 11 with inner rib ring in actual fabrication, the manufacturing cost and procedure will be increased and the total weight will be increased with about 40t when some rigidity support components cannot be removed. It is desired to propose an advanced H. Zhou et al. Ocean Engineering 196 (2020) 106812 technique not only to control the welding deformation but also to decrease the fabrication cost and weight of cylindrical leg structure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure15-1.png",
        "caption": "Fig. 15. Coupling with Rim Pumping",
        "texts": [
            " It was quite unsuitable for high powers, or for working continuously in a partially filled condition, owing to the low head available for circulating through a cooler. Furthermore, because a fixed casing was necessary to enclose the rotating parts, the rate of heat dissipation was very low, and on these several grounds the design was abandoned. An example of the rim pumping idea was carried out in connexion 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from with a sump tank only, and no gravity tank, in the fluid coupling shown i n Fig. 15 and Fig. 18, Plate 4. The coupling is of the conventional Vulcan type completely enclosed in a fixed casing, in which an annular channel surrounding the rim is formed. The lower part of the casing is extended to form a sump tank with bearings supporting the impeller casing and the runner. A pumping disk or fin, mounted on the rotating casing, entrains oil within the annular channel of the fixed casing, which it delivers together with oil discharged by the usual leakoff nozzles from the working circuit into a port in the top half of the unit"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002072_j.simpat.2020.102236-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002072_j.simpat.2020.102236-Figure4-1.png",
        "caption": "Fig. 4. Mesh of fluid volume.",
        "texts": [
            " The equation of shaft radial micromotions is defined as: F\u2192ex + F\u2192sup ( et \u03be, e t \u03b7, e\u0307 t \u03be, e\u0307 t \u03b7 ) = M e\u0308\u2192 (35) In the current section, the results of the extensive sets of simulations for the HJB-EGP system are proposed. The specifications of the components are detailed in Table 1. Unless specified otherwise, these parameters are implemented in all simulations. Different types of mesh models are proposed for the EGP displacement chamber and HJB lubrication gap before simulation. A specialized tool within PumpLinx is used to generate a computing grid, as shown in Fig. 4(a). A high-quality mesh is found at micronlevel clearances in the EGP displacement chamber, which plays an important role in the simulation in order to provide precise results. Furthermore, the bearing calculation domain is discretized into a number of discrete grids, as shown in Fig. 4(b). The terms N, S, W, and E represent the primary node, and n, s, w, and e represent the secondary node. To exclude the grid influence, a sensitivity analysis was performed at 2400 r/min and 3.8 MPa. Several mesh models of the HJB and EGP are presented in Tables 2 and 3. Note that the EGP mesh models have the same cell number at the inlet and outlet ports but different cell numbers at the gear zone. The flow rates in relation to the number of elements are shown in Fig. 5(a). The flow rates obtained by the a2 and a3 mesh models present similar values"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000678_978-3-319-45781-9_19-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000678_978-3-319-45781-9_19-Figure3-1.png",
        "caption": "Fig. 3. Test parts: existing part and final part.",
        "texts": [
            " In machining, the accessibility of cutting tools is one of the major constraints to be taken into account during the identification and the extraction of manufacturing features. If the build of an AM feature cause an inaccessibility of cutting tool in the next machining operation, it has to be built after the machining operation. The constraints of part clamping in machining stages should also be considered. The proposed feature extraction process contains five major steps as shown in figure 2. The proposed extraction process is demonstrated using the case study presented in figure 3. For this purpose, all steps were performed manually using a CAD software. The pocket (P), the hole (H) and the surfaces (fS1 to fS7) of the final part require a high surface precision. The roughness of the surfaces (eS1, eS2, and eS3) of existing part satisfies the quality of the final surfaces (fS1, fS2 and fS3). The steps of process are outlined as follows: Local coordinate system definition and Positioning: The first step consists of defining a local coordinate system for each CAD model of existing part and final part",
            " Thus, a sufficient over-thickness for the finishing stages leaving on these surfaces is taken into account in generating the CAD model of AM features (e.g., AMF_2). It is estimated as a function of roughness of surfaces generated by AM process, the required quality of final surfaces, and the surface quality achieving by machining. The over-thickness will become the rough state attribute of machining features after AM process. The machining features after AM process, denoted as MFa, are determined from the functional features of final part (for example, the functional surfaces fS4 to fS7, the holes (H), and the pocket (P) in figure 3). The rough state attributes of MFa features defined by the over-thickness integrated in AM features, or a plain material state (particularly in the case of drilling holes). In figure 5b, the machining features, MFa_{1, 2, 3, 4, 6}, correspond to functional surfaces of final part; and MFa_5 corresponds to the hole feature (H). The research focused on the feature extraction process in a context of remanufacturing. The proposed approach allows an effective extraction of both MFs and AMFs features from the CAD model of existing part and final part, and the knowledge about constraints of AM process and machining process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000763_s10443-016-9565-5-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000763_s10443-016-9565-5-Figure14-1.png",
        "caption": "Fig. 14 Comparison of the composite damage in experimental and finite element simulation under the axial (warp direction) compressive loading",
        "texts": [
            " Figures 12 and 13 display the detailed damage developments of 1 RVC and 2 RVCs under transverse compression in the weft direction. Unlike the results of axial compression in the warp direction, the damage initial, damage evolution, final failure and damage distribution and location damage of 2 RVCs are almost the same with that of 1 RVC. At the max loading point, the damage appears only the weft area next to the blinder warp. Afterward, the damage becomes serious, but there are no widely spread. Moreover, the max stress points of 2 RVCs occurs early at \u03b5 = 0.66% in comparison with that of 1 RVC. Figure 14 shows a comparative analysis of the numerical and experimental damage morphologies under axial (warp direction) loading. The fractured yarn plane is most consistent with the numerical one. Especically, noted that the failure shape of 2 RVCs is more close to that of experiment. It can be noticed that the damage modes of the axial compression are totally dominated by the binder warp failure, matrix crack and warp fracture. Moreover, The blinder warp exhibits a more dispersed damage, with damage occurring at different position in each blinder warp owing to the greater amount of yarn distortion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003630_047134608x.w1111-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003630_047134608x.w1111-Figure3-1.png",
        "caption": "Fig. 3. The Euler angles.",
        "texts": [
            " In order to be able to prescribe a desired trajectory for the aircraft center of mass, it is necessary to determine the position and orientation of the aircraft (or equivalently its body-fixed axis system) relative to the earth-fixed reference frame or axis system. Accomplishing this requires the introduction and definition of three angles referred to as the Euler angles. These angles can be used to uniquely prescribe the orientation of the body-fixed axes relative to the earth-fixed axes through three ordered rotations as shown in Fig. 3. Note that the order of the rotations , , is essential to the definition of the Euler angles. Equations Expressed in Body-Fixed Coordinates. The vectors F, vc, G, and h can be conveniently expressed as components in the body-fixed xyz axes of Fig. 1 as follows: where i,j,k represent unit vectors parallel to the x, y, and z body-fixed axes and where X, Y, and Z represent the components of the aerodynamic and propulsive forces alone, with the contribution of the gravitational forces now included in such terms as \u2212mg sin "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002152_978-981-13-3549-5-Figure5.16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002152_978-981-13-3549-5-Figure5.16-1.png",
        "caption": "Fig. 5.16 Snapshots of rover simulation in RoSTDyn [96]",
        "texts": [
            " Based on terramechanics, dynamics and kinematics, the high-fidelity simulation can support 3D predictive display for direct tele-operation of lunar rovers. If the scientific \u201cgoals\u201d are sent to the rovers directly, the virtual rover will switch to autonomous navigation (AutoNav) mode and move forward based on virtual stereo vision; if \u201cpaths\u201d such as line, arc and steering-in-place are sent to the rover, the locomotion mode will be booted and drive the motors directly by setting their \u201cpositions.\u201d The simulation snapshots of China\u2019s Yutu lunar rover in RoSTDyn are shown in Fig. 5.16. The interactive virtual planetary rover environment program helps engineers with direct tele-operation of lunar rovers. It is realized by comparing the feedback motion of an actual lunar rover with the counterpart of the virtual lunar rover predicted by 128 5 Integration and Scheduling of Core Modules RoSTDyn. The virtual motion is ahead of the actual motion, and there is a twofold time delay between them. The time delay could be modified by setting an additional one before the command sequences are sent or adding one to the virtual rover"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002406_s42242-020-00083-7-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002406_s42242-020-00083-7-Figure1-1.png",
        "caption": "Fig. 1 Three-dimensional virtual prototype of CPM",
        "texts": [
            " Section 2 designs and manufactures CPM and constructs the dynamic model. Section 3 completes the controller design and stability analysis. Sections 4 and 5 present a computer simulation and hardware-in-loop simulation (HILS) experiment, respectively. Section 6 conducts a cell puncture experiment on zebrafish embryos and, finally, summarizes the paper. A bridge-type displacement amplification mechanism (BTD /AM) is designed based on the compliant mechanism principle, which does not bring driving clearance to CPM, thereby ensuring a high-precision motion effect. Figure 1 shows the 3D virtual prototype of CPM. The two ends of PEAs are separately fixed to the BTDAM by a locking and preloading mechanism. This scheme eliminates the connection gap and applies a certain preload to the PEAs. The centroid of the injecting pipette is locked to the BTDAM, whereas the tail is slidably connected to the guiding hole to ensure that the injecting pipette strictly moves along a straight line. When PEAs are extended, the flexure hinge of BTDAM is elastically deformed. This condition drives the injection pipette to output a large displacement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003626_piae_proc_1922_017_035_02-Figure21-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003626_piae_proc_1922_017_035_02-Figure21-1.png",
        "caption": "FIG. 21. F I G . 2 2 . .",
        "texts": [
            " Here, it shoukd r 7 Jhis ' usually consists of one or more helioal springs in series with the main laminated springs, and it has the dynnniical effeoti of introducing flexibility or elastiaity without friction. I n the chlaracteristic diagram it is easily represented because the load on tho helioal or auxiliary spring is praotimlly the mme as that on t,ho main or laminated springs, and thus the deflections of the * See Proc. 1n.t. C. E., Vol. CLXXSV., Part 111. at The University of Auckland Library on June 5, 2016pau.sagepub.comDownloaded from PRINCIF\u2019LES OF VEHICLE SUSPENSION. r305 tuo springs are added together to give the total deHectioii of the sjsteiu, see Fig. 21 and conipare it with Fig. 18. A further variation, and the last that we need oonsider before prooeding to generalisation, is the so-called shock-absorber, which as a rule memly adds fluid friction to the system. This is characterised by adding the diagrams in Figs. 17 and 18, see Fig 22. and it is obvious at a glancle that shook-absorbers of this type aro useless on springs which already h a w much solid friction. Shock-absorbers of the fluid-friction type are useful as remedies OIL mrs where the springs have iiisufficient pip or few leaves, i4 which case they inay approach Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000489_2016-01-1560-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000489_2016-01-1560-Figure5-1.png",
        "caption": "Figure 5. Lateral deformation of the screw predicted by the Solidworks model.",
        "texts": [
            " As seen from the figure, the load distributions predicted by the Solidworks model for the L and U contact points are in good agreement with those of our proposed model; both models show that the contact load distributions of the balls are not uniform for the L and U contact points. However, Mei et al.\u2019s model predicts uniform load distributions on all balls, which is inaccurate. The weakness of Mei et al.\u2019s model lies in the fact that it cannot capture the coupling between axial, torsional and lateral deformations [14,15,16] created by the oversized balls. This fact can be seen by examining the lateral deformation of the screw predicted by the Solidworks model as shown in Figure 5, which clearly shows the deformation of screw in both x and y directions solely as a result of the oversized ball preload. Figure 6 compares the centerline displacements of our proposed model and the Solidworks model; both models are again in good agreement in predicting the lateral deformation of the screw. The maximum lateral deformation magnitudes in the x-z and y-z planes are around 8 \u03bcm and 35 \u03bcm, respectively. Such magnitudes of lateral deformation are too significant to be neglected in modeling the contact load distribution of the balls"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000931_s00170-019-03553-0-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000931_s00170-019-03553-0-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of the deposition setup during a thin wall deposition",
        "texts": [
            " However, this is only feasible if the laser metal deposition can produce good-quality homogenous material over a wide range of blended compositions. This study aims at investigating this feasibility. The deposition of the materials under this study was carried out on a system equipped with a CNC stage and an edge nozzle feed setup connected to a Bay State Technologies Model 1200 powder feeder. A 1-kW fiber laser from IPG Photonics was used as the power source. A schematic diagram of the deposition setup is shown in Fig. 1. The powders employed in the study were elemental copper (99.9% pure, from Oerlikon Metco, \u2212 100/+ 325 mesh, gas atomized), elemental nickel (99.9% pure, from Atlantic Equipment Engineers, \u2212 100/+ 325 mesh, gas atomized), and Delero-22 alloy (composition in Table 1, from Kennametal Stellite, \u2212 100/+ 325, gas atomized). The powder feedstocks were prepared by mixing precisely weighed amounts of each powder in a container using a Turbula mixer. Each blend of powder was mixed for 30 min to ensure uniformity within the blend"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003542_pime_proc_1948_158_045_02-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003542_pime_proc_1948_158_045_02-Figure16-1.png",
        "caption": "Fig. 16. Diagram of Vectors",
        "texts": [
            " putting A / 2 c = a, Alternatively, a small groove might be turned at the root of the h g e of the present 1 in 20 profile, so as to remove in advance the offending part of the worn profile. But there may be practical objections to this, and in any case the effect on the worn rail profile would have to be carefully investigated. Acknowledgements. The experimental work was carried out in the Engineering Laboratory of the University of Cambridge, with help and financial support from the London, Midland Instrument Company for advice about recording on celluloid. b c 2sx1 p+y -sin ry - B$ -h cos 1 <a - * (4) The left-hand side of (4) gives the vector diagram in Fig.16. For no interference by the jl and y plays throughout the cycle, the dotted vector PQ must be less than (B+y)/a, and the point Q must therefore lie within a circle of centre P and radius ( P + y ) / a ; whence, by trigonometry from Fig. 16, and Scottish Railway Company. The authors wish to express sin B+J(q)Z-(i-cosB)z>- 4sc their thanks to that Company, and also to the Cambridge h > sinB-J(\u2019+)\u2019-(l-cosB)z . . . (5) A P P E N D I X I Condition (lb) gives a diagram which is a mirror image of Substituting from (3) in (2) we get Fig. 16, and therefore repeats condition (5). Let the yawing play of each axle relative to the bogie frame be & p ; it is due to the fore-and-aft play of the axleboxes in their guides. Let the lateral play of each axle relative to the frame be & d ; it is made up of the lateral play of the axleboxes Let the wheelbase be 2c, and let d/c = y. in their guides and the end play of the journals in their brasses. - 477c Which &eS, as another condition for no interference by the B Plays thoughout Assume that in their mid-positions the axles are parallel and in track"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001269_tie.2019.2950838-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001269_tie.2019.2950838-Figure10-1.png",
        "caption": "Fig. 10. Fabricated prototypes. (a) Prototype A: without an isolating mechanism. (b) Prototype B: with an isolating mechanism.",
        "texts": [
            " The electrical signal is generated by the function/arbitrary waveform generator (RIGOL DG1022U) and then sent to a power amplifier (HSA4051, NF Corporation) before driving the two prototypes. The voltage output by the amplifier is measured as 40 V with a 20 V bias. A noncontact laser sensor (OPTO NCDT2300-2) with a measuring range of 2 mm and resolution of 0.03 \ud835\udf07m is used to measure the displacement. Its maximum sampling frequency is 49.02 kHz. The output data from the laser sensor are recorded by ILD 2300 software and processed on a PC terminal. Fig. 10 shows the fabricated prototypes. A steel cylinder and a carbon fiber rod are used as the load and the light slider, respectively. A copper block serves as the inertial body. Both the slider and the inertial body are glued to the ends of the piezo stack by a type of epoxy resin (DP460). The preload force F (>10 N) provided by the compressed spring is considerably larger than the gravity of each prototype. Thus, structure imbalance can be ignored. Regardless, to weaken the influence of imbalance, the stick-slip motion always occurs in the left area of the slider"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000130_coase.2019.8843076-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000130_coase.2019.8843076-Figure7-1.png",
        "caption": "Figure 7: Path planning",
        "texts": [
            " In order to evaluate the control strategies, a test cube with an edge length of 40 mm has been welded. The tests were welded with the following parameters and boundary conditions: wire feed: 4 m/min; welding speed: 0.6 m/min; process modification: controlled short circuiting arc; filler material: 1.2 mm EN ISO 14341-A: G 3Si1; substrate plate: 3 mm S235JR; shielding gas: Ar: 98%, CO2: 2%; interpass temperature: 100 \u00b0C \u00b1 10 \u00b0C. The CAD model and the corresponding path planning for 27 layers without top and bottom layers, 50% infill and one wall layer are visible in Figure 6 and Figure 7. The following disturbance variables were deliberately provoked: The layer height and the bead width were chosen lower in the path planning than to be expected for the given parameter set. The actual volume build-up will thus exceed the parameterized layer height and the total expected volume. => ? Q\\^ Q> _`{\\ Q| } > Q` | > _`{\\ Q| } > } | > `\\^> the dimensions 10x10x40mm was connected to the cuboid. At this point, due to the incorrect parameterization of the bead height and width, larger material accumulations will occur especially on the disturbance cube"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure3-1.png",
        "caption": "Fig. 3 Schematic diagram of plane rake face",
        "texts": [
            " As the basis of mathematical derivation, the cutter coordinate system S2 (o2, x2, y2, z2) is first established as shown in Fig. 2, where axis z2 coincides with the centerline of the cutter. On the premise that the cutting edge is error-free, the rake face can be constructed in two ways: by a plane that is easy to manufacture or by a curved surface that can ensure consistent and reasonable working rake angle. The plane rake face is constructed in the following way to form rake angle. As shown in Fig. 3, vectors lbz, lbt, and lb are defined by the criterion of passing through the midpoint D (xD, yD, zD) of the addendum of the conjugate surface (see literature [6]). Vector lbz is the unit vector along the direction of axis z2. Vector lbt directs along the rotation direction about axis z2. Vector lb is located in the plane defined by lbz and lbt and has an intersection angle (which is equal to the helical angle \u03b22 of the cutter) with lbz. PointDz is the projection of pointD on axis z2. Through pointDz, line segmentDzQo is constructed along the direction of vector lb, and its length is R0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001905_j.mechmachtheory.2020.104095-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001905_j.mechmachtheory.2020.104095-Figure6-1.png",
        "caption": "Fig. 6. Example virtual system (a). Initial status. (b). After bending. In Fig. 6 , the dotted line indicates flexbile parts in example virtual system, the rigid line indicates rigid parts in example virtual system.",
        "texts": [
            " 5 (b), after electrified, the moving platform can rotate and slide slightly around the mass block. The stretch of central chain can be reduced. That makes the axial deformation of the central chain and cantilever beam negligible in dynamic modeling. The pose of the central chain and the cantilever beam can be described by rotation angles. Besides, the half-cylinder of mass block makes it easier to obtain the constraint relationship between central chain and kinematic chains. For illustration purposes, we use the virtual system shown in Fig. 6 as an example. The virtual system consists of two rigid beams, one flexible beam, two mass blocks, and four flexible steel plates. Beam h 1 is fixed at the center of beam v 1 . Beam h 1 is vertical to beam v 1 . Beam v 1 is connected with flexible beam v 2 . Two mass blocks are tangent to beam h 1 . Each mass block is fixed at the top of two serial flexible steel plates. In Fig. 6 , s 3 is the center of circle 1, s 6 is the center of circle 2. Circle 1 and circle 2 are tangent to beam h 1 . s 2 and s 5 are the point of tangency. The center of the bottom line of rectangle 1 and rectangle 2 are s 4 and s 7, respectively. The initial status of the virtual system is shown in Fig. 6 (a). Then, we bend flexible steel plates, thereby bending beam v 2 and rotating beam v 1 . After bending, the virtual system is shown in Fig. 6 (b). Since circle 1 and circle 2 are tangent to beam h 1 , we can get s 3 s 2 \u22a5 h 1 and s 6 s 5 \u22a5 h 1 . That means s 3 s 2 \u2016 s 6 s 5 . Taking the fact that s 3 s 2 = s 6 s 5 and s 3 s 2 \u22a5 s 2 s 5 into consideration, we can say the square s 3 s 2 s 5 s 6 is a rectangle. That makes s 3 s 6 \u2016 s 2 s 5 and s 3 s 6 = s 2 s 5 . In real work, the rotation angle of the mass block is not large. Values of s 3 \u2019s 4 s 3 and s 6 \u2019 s 7 s 6 are less than 0.05 \u00b0. Thus, we assume s 3 \u2019s 4 s 3 = s 6 \u2019 s 7 s 6 \u22480 \u00b0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003859_pime_proc_1995_209_421_02-Figure33-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003859_pime_proc_1995_209_421_02-Figure33-1.png",
        "caption": "Fig. 33 Diagram depicting the ball wear pattern from ballondisc tests (60) at a 20 per cent slide-roll for a lubricant seeded with 0.5-1.0 pm diamond particles",
        "texts": [
            " This is thought to be largely due Q lMechE 1995 DEBRIS AND ROUGHNESS IN MACHINE ELEMENT CONTACTS 169 --- I rn I 0 I 2 Number of revolutions million rev Wear loss for two debris particle materials at a concentration of 1 g/l of lubricant Fig. 31 to a transition in wear mode from an embedding/ abrasive mode with the higher fracture toughness materials to a turnblindabrasive mode with the lower K,, debris. This latter case occurs when the particle fragmentation is suficient to allow particles to tumble through the contact. Dwyer-Joyce (60) performed further work on these mechanisms using a ball-on-disc machine where the degree of rolling and sliding could be controlled. Figure 33 shows a schematic representation of many of his results for small, high-fracture toughness particles and for low-fracture toughness particles. The transition from tumbling to abrasive ploughing seemed to be related to the ratio of particle size entering the contact to the thickness of the contact a m . The transition from tumbling to grooving has also been reported by Williams and Hyncica (64) who worked with the more conforming, and consequently thicker, hydrodynamic a m geometry offered by a flexible metal belt wrapped around a rotating shaft. With this geometry they suggested a ratio of particle size to film thickness of about 2 as the transition, which the present results on much thinner EHD films more or less confirm. For example, disc machine tests with diamond abrasive particles from 2 pm and above were seen to change the picture of Fig. 33 to show abrasive grooving across the whole contact path. Q IMcchE 1995 Prw Inatn Mech Engrs Vol 209 170 R S SAYLES 4 SUMMARY A N D IMPLICATIONS FOR FUTURE RESEARCH The relationship between roughness and A ratio for two-body tribological contacts has been reviewed and some areas of future work, particularly in relation to defining relevant roughness values in relation to elastic conformity and the influence of asymmetric roughness structures, should be addressed in future research. These effects and others should also be examined in relation to the modes of contact outlined from boundary lubrication work (53)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000132_b978-0-12-814245-5.00035-9-Figure35.15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000132_b978-0-12-814245-5.00035-9-Figure35.15-1.png",
        "caption": "FIGURE 35.15 Parameters defining robot pose: (A) robot with mounted RONNA stereo; (B) robot with a mounted surgical instrument which uses the same virtual TCP as RONNAstereo. TCP, Tool center point. Reproduced with permission from S\u030culigoj F, Jerbic\u0301 B, S\u030cekoranja B, Vidakovic\u0301 J, S\u030cvaco M. Influence of the localization strategy on the accuracy of a neurosurgical robot system. Trans FAMENA 2018;42:27 38. doi:10.21278/TOF.42203.",
        "texts": [
            " Localization strategies are defined through robot approach angles, orientations, and movements during fiducial marker localization in physical space, as well as motions for positioning onto target points [53]. Robots with six or more DoF can approach targets in their workspace with an unlimited number of different orientations resulting in different joint configurations. Furthermore, each trajectory can be reached by the robot in numerous configurations, that is, orientations around the longitudinal tool axis [50], as shown in Fig. 35.15. With the ability to approach fiducial markers from 3 5 . R O N N A G 4 R o b o tic S yste m different angles during localization and the ability to change the angle around the x-axis while positioned in the targeted trajectory, RONNA can implement different robot localization strategies. The goal of robot localization strategies is to decrease errors in patient registration and instrument positioning with respect to their designated target pose. Each strategy uses a set of F which contains n localization poses F5 F1; . . . ;Fnf g and a set T which contains m target poses T 5 T1; . . . ;Tmf g. As shown in Fig. 35.15A, the position of the robot end effector flange is determined by six joint states \u03b8. Connected to the end-effector flange is the tool, whose end pose relative to end-effector flange is fixed and defined using a Cartesian coordinate system translation vector (x, y, z) and a series of three rotations (\u03b1, \u03b2, \u03b3) according to the Z Y X convention. Combining the transformation from the robot base to the end-effector flange with the tool transformation, the robot tool pose is given as the transformation R-TCP (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001811_s12008-020-00670-z-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001811_s12008-020-00670-z-Figure3-1.png",
        "caption": "Fig. 3 Example of the information that students can identify in a 4-DoF manipulator using the platform",
        "texts": [
            " Calculate the homogeneous transformation matrix for each pair of neighboring joints. 4. Multiply the homogeneous transformation matrices obtained in the previous step to obtain the position of each joint and the final effector with respect to the manipulator\u2019s base. 1 3 To perform these steps, students with the help of the platform must identify each movement of the robot, the number of links, the types of joints (prismatic or revolute), and the direction of the movement of each joint to define its axis. Figure\u00a03 summarizes the information that students can identify using the virtual model of the robot. Figures\u00a02 and 3 show the different blocks and elements of a full 4-DoF serial manipulator. In Fig.\u00a02, the \u201cForward kinematic section\u201d has inside two functions (also called blocks) that calculate the forward kinematics of the robot; i.e., this section calculates the translation and rotation transformation matrices from each joint using the DH parameters. The first block is \u201cHomogenous Transformation Reference frames,\u201d in which students must set the reference frames of each joint",
            " In this step, users must select the correct blocks and connections among them. The student must select between the \u201crotation variable axis\u201d and the \u201ctranslation variable axis\u201d blocks according to the movements of the robot, and they must delete those that are not necessary. These blocks represent the four transformations between reference systems established by the DH method. (Students must select only four blocks that define the DH parameters.) This process must be repeated for all the Transformation blocks. For example, in Fig.\u00a03, the joints 1, 2, and 4 perform angular movements; then the student must select the \u201crotation variable axis z block\u201d inside the Fig. 5 Blocks inside the Transformation A1, A2, and A4 blocks Fig. 6 Blocks inside the Transformation A3 block 1 3 transformation blocks A1, A2, and A4 (see Fig.\u00a04); the joint 3 performs a translational movement, so the \u201ctranslational variables axis z block\u201d must be selected inside the transformation block A3 (see Fig.\u00a06). The platform allows the user to improve their skills to position the reference systems in each joint of the robot (an example of a 4-DoF robot with the reference systems is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001880_j.mfglet.2020.08.004-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001880_j.mfglet.2020.08.004-Figure1-1.png",
        "caption": "Fig. 1. Contrasting the workflows in a traditional job-shop vs. an SMM to realize a specified geometric shape, microstructure and morphology (a) Hybrid manufacturing pathway with printing and machining to create (i) net geometric shape, laser assisted heat treatment to vary (ii) microstructure and a diamond grinding pin and Dremel tools to change (iii) morphology (b) Conventional manufacturing process chain to realize the same output using five different machines/equipment like a sand cast for consolidation of material, milling machine to create a rough finish, heat treatment furnaces to vary microstructure and grinding and polishing machines to impart final mirror-like finish.",
        "texts": [
            "\u2019s [11] metamorphic manufacturing is one such concept which includes metal forming techniques (\u2018\u2018robotic blacksmith\u201d) to modify the shape and grain structure in incremental steps. We conceptualize a smart manufacturing multiplex (SMM) that builds on these efforts as follows. The current concept of SMM is implemented on a hybrid machine tool LENS MTS 500 from Optomec. It consists of a directed energy deposition (DED) laser head for printing near-net shape components and a vertical rotating spindle to perform machining, grinding and polishing processes (see Fig. 1). This implementation allows the control of not just the geometric dimensions, but also of the morphology (via grinding and polishing), and, to a limited extent, the material composition and microstructure by adjusting the laser heat-treatment parameters, all using the same datum in a machine tool. The infusion of sensor technologies and data science/AI algorithms transforms HM into a smart manufacturing multiplex (SMM) which can not only execute process chains that traverse the material and process parameter space, but also enable tracking across various processes",
            " This section presents the architecture of an SMM and an initial demonstration of the following capabilities: (1) execution of multiple process chains to improve productivity, (2) integration of a sensor wrapper into the process workflow to track process state and enhance quality assurance, and (3) manipulation of energy input to an HM machine tool to reduce machining forces. A summary of the implementation challenges and limitations of the SMM and a note on future directions to continue the present work is discussed. For the current SMM implementation, Optomec MTS 500 hybrid machine tool with the capability for metal powder deposition and a machining spindle was used. Fig. 1 contrasts the manufacturing workflow of a custom part on the SMM (Fig. 1(a)) versus in a traditional job-shop (Fig. 1(b)). Traditionally, a sample or a workpiece from a near-net shaping process has to be removed from the platen on which it was fabricated and is followed by machining to meet the geometric (tolerance) specifications. The part is subsequently heat treated to impart the desired microstructure and polished to control the surface finish and morphology. Each individual step is usually executed on a separate machine tool, and it incurs significant time to setup tooling and fixtures, to determine the datum planes for measurements, as well as to transport between machines"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001019_rpj-07-2018-0182-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001019_rpj-07-2018-0182-Figure10-1.png",
        "caption": "Figure 10 Sectional view in FEA software",
        "texts": [
            " The top die surface is extracted from the solid CAD model and considered as analytical rigid surface. The tube was modeled as a deformable surface of 0.3 mm thickness based on elastic and plastic material properties of stainless steel SS304 as shown inTable III. The tube is meshed with 2424 elements and 2506 nodes of S4R four-noded, quadrilateral, shell element with reduced integration and a large-strain formulation. Bottom die is meshed with 36303 linear tetrahedral elements of type C3D4. Material property of ABS-P430 obtained from flexural test is assigned to bottom die. Figure 10 shows assembled view of tube with top and bottom die in Abaqus CAE software. The deformation of the tube occurs only in the bulging and folding steps, while the other steps do not involve any shape change. Therefore, only bulging and die closing steps were considered for the simulation. In both steps, the bottom die is fixed by applying encastred condition. A frictional contact defined with coefficient of 0.25 (Leacock et al., 2015) is between the tube surfaces and die surfaces. Axis-symmetric constrains were applied at the longitudinal section edges of the tube"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003768_1.2833790-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003768_1.2833790-Figure3-1.png",
        "caption": "Fig. 3 Coordinate transformation for a herringbone bearing",
        "texts": [
            "org/ on 02/05/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use \\2H rde 2 M, = - ph^ dp Tlfl'dr (1) (2) For SPB, TPB, and TFB, the Reynolds equation in the form of mass flow conservation can be written out and solved directly in the (r, 6, z) coordinates system. However, for HB, as the groove and the land are formed by spiral lines, nonorthogonal boundary-fitted coordinates system (^, rj) is used on the inner side (r, s r ^ r\u201e) and (^', 77') on the outer side (r^ ^ r s r\u201e) of the bearing's surface as shown in Fig. 3. In this figure, the boundary lines of the grooves and the lands are logarithmic spiral lines and /3 is a spiral angle. The transformation (James and Potter, 1967) between the (r, 9) and the (^, rj) systems is (3) (4) 8 dr r^ = r(e-f(r)) d 1 + ln(C/r,.) d dC, tan /3 drj d _ d rde \" dri (5) (6) In this transformation, a couple of groove and land in the (r, 6) system as shown in Fig. 3(a) is transformed into the one in the (^, 77) system as shown in Fig. 3(fo). Assuming the mass flow normally passing through unit length in the ^ and 77 direc tions be M; and M,, respectively, M^ and M^ can be derived through the transformation as follows. Mr r^Mr'\u2014Mt or) art dr Va (7) M,^[-rf^M.^f^M,]/ry (8) where. 7 = 1 + a = 1 (1 +ln(^/r , .))^ tan^/3 (9) (10) Assuming an isothermal and viscous laminar flow, the Reyn olds equation in the form of mass flow conservation with respect to the control volume that is shown in Fig. 4 becomes: AC._, , , , , AC; + (M,),+,,"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001968_j.icheatmasstransfer.2020.104868-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001968_j.icheatmasstransfer.2020.104868-Figure8-1.png",
        "caption": "Fig. 8. Schematic view of: (a) standard design of assembly with matched radiation baffle (ARB), (b) proposal of design of inner radiation baffles (IRBs 1\u20133) to production of large IGT blades using Bridgman method. Ceramic shell mold is not shown in figures.",
        "texts": [
            " In those investigations [19], the baffle opening was limited by the maximum extension of cross-section, mainly of the platforms. Complete shielding between the heater and cooling area resulted in a further increase of temperature gradient and a significant refinement of microstructure to value PDAS = 380 \u03bcm. This comparison of various baffle geometries shows that the elimination of the unfavorable gap around the airfoil and the central rod gives the possibility to further microstructure refinement in the blades manufactured using the industrial-scale Bridgman process. Fig. 8a shows a schematic view of the design of assembly and external radiation baffle that can be used to produce large IGT blades with columnar grain structure using the Bridgman method. In this example, the external radiation baffle is mainly matched to the shape of root and platform of four blades and the gating system as much as possible. However, for the large IGT blades, the gap between the baffle opening and blade is wider compared with that, when the small aero engine turbine blades are produced",
            " Therefore, it seems that the improvement of the baffle matching should be important, especially during the production of large IGT blades, in which the width of gap increases with increasing dimensions of the root and platforms relative to the thickness of blade airfoil. However, a greater reduction of the analyzed gap and the improvement of temperature gradient in IGT blade is only possible by using the inner radiation baffle technique in the Bridgman method. On the basis of analysis of the obtained research results and those presented in references [19], the new design and a proposal of the implementation of IRBs technique for the production of large IGT blades using Bridgman method has been discussed in this article. Fig. 8b shows the shape and location of IRBs in the assembly, taking into account the same geometry of external radiation baffle as it used to be in the standard process (Fig. 8a). The geometry of IRBs and method of their location in the assembly differ from those shown in Fig. 2 for the production of small aero engine turbine blades. The IRBs were divided into sections (designated as IRBs 1\u20133) in order to match them more precisely to the shape of the blade (IRBs 1 and IRBs 2) and to position them easily in the assembly (IRBs 3) (Fig. 8b). The presented solution provides the reduction of the gap width by the IRBs, mainly at the airfoil height, simultaneously keeping the area of ARB as large as possible. However, the opening contour of external radiation baffle is sufficiently matched to the geometry of the starter and root. Therefore, the use of IRBs at the height of starter and root is not technologically justified. Attaching the IRBs to the mold seems complicated, but in some cast processes, this technique may be the only one to increase the solidification parameters above the critical values to prevent the defect formation and improve the castability of directionally solidified large IGT blades [19,23]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002221_0278364919897134-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002221_0278364919897134-Figure8-1.png",
        "caption": "Fig. 8. CAD model of membrane mold base for spherical configuration fiber layup (a) and completed fiber-reinforcement mesh (b). The outermost circular fiber is included to provide extra stiffness in the completed membrane at the clamping surface.",
        "texts": [
            " The increase in stiffness prevents rupture of the buoyancy bladder at increased depths and also helps maintain the desired shape even under significant external loading. Fabrication of the membranes tested in this study was conducted in two main steps: fiber layup (Section 3.1) and two-stage elastomer molding (Section 3.2). The fiber-reinforcement patterns for the membranes designed for this study were prepared on 3D-printed molds (3D Systems ProJet MJP 2500 Plus). The mold designed for the spherical final configuration is shown in Figure 8. Extruded posts trace the path of each fiber, whose orientation was determined based on the modeling in Section 2. The maximum fiber density is determined by the ability to lay out these posts without them crossing the path of another fiber. Small extrusions jut out from each post to raise the fiber off the base of the mold, ensuring that elastomer will encase as much of the fiber as possible during the molding process. If the fibers make contact with the base of the mold, it is likely that they will delaminate while removing the membrane from the mold, ruining the sample in the process. The larger posts on the outer edge result in holes for clamping screws and serve as an anchor point for each of the radial fibers. Cotton fibers were chosen to enhance bonding with an elastomer matrix. Ecoflex 30 elastomer (Smooth-On, Inc.) was chosen as the matrix of the composite owing to its combination of large strain at rupture (allowing for large deformations) and stiffness (to prevent issues with clamping). The fibers are arranged as seen in Figure 8, with circular fibers shown in orange and radial fibers shown in blue. An additional, outermost circular fiber lies under the clamp to provide extra, more uniform stiffness for a proper seal with the clamp base (see Section 4.1 and Figure 8). While the radial fibers have anchor points to aid in the layup process, the circular fibers do not have anything to hold them in place. Therefore, they are laid down first so the radial fibers can hold them in place once the layup is completed. Each circular fiber was pre-tied with a simple noose knot before being placed on the mold base (Figure 9(b)) and tightened around its respective guide posts. The two ends of the fiber were then tied with a square knot before being trimmed as short as possible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003103_tee.23458-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003103_tee.23458-Figure11-1.png",
        "caption": "Fig. 11. SRM experimental test setup: (a) Controller (b) SRM",
        "texts": [
            " The torque ripple of MPTC technique is always kept at a lower level than DITC before and after increasing or reducing the load torque, as shown in Fig. 10. At the same time, it can be shown that the three-phase torque distribution of MPTC more intelligent than DITC. Moreover, after the value of load torque is decreased from 10 to 2N\u00b7m, the peak current of DITC is 3.81 A, while the peak current of MPTC is 3.14 A. In summary, simulation results prove that MPTC control method has lower torque ripple and efficiency in the situation of load mutation. 6. Experimental Results The test setup is decipited in Fig. 11, the controller and SRM are illustrated in Fig. 11(a) and (b), respectively. Specifications of some devices are listed in Table III. In the experiment, the sampling period of the proposed method and DITC is set to 100 us, and the given speed is set as 500 r/min with 5 N\u00b7m load torque, the experimental results of MPTC and DITC methods are decipited in Fig. 12(a) and (b), respectively. From the torque waveforms, results show that the Tpeak of MPTC method is reduced by 1.07 N\u00b7m compared with DITC at same experimental conditions. From the current waveforms, it can be noticed that the peak current of the MPTC is smaller and smoother 6 IEEJ Trans (2021) than DITC, whereby the proposed MPTC method copper losses is lower than DITC"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000728_icelmach.2016.7732503-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000728_icelmach.2016.7732503-Figure3-1.png",
        "caption": "Fig. 3. Stator and rotor geometry of the modeled SynRM.",
        "texts": [
            " \u03be (B) = kp ( khB2f + keB2f 2 ) [ W/m3 ] (7) where kh and ke are Hysteresis and Eddy current coefficients which depend of the magnetic material, f is the electric frequency, and kp is a factor related with the lamination punching. Even though, this approach is essentially empirical, it is well suited for fast and rough iron loss determination since can be easily integrated in Finite Element simulation [19], [20]. The electrical machine used in this research consists of an industrial-type SynRM available in the market. The manufacturer sells this machine with its own frequency converter for motor operation in applications such as pumping, fans, conveyors, mixers, etc. Fig. 3 shows the structure of the stator and rotor of the machine and Table I shows the main parameters obtained from the name plate, disassembling and experimental tests. The program used for modeling of the SynRM is FEMM (Finite Element Method magnetic) 4.2. In order to make the study more efficient, the analysis is reduced to a fourth of the whole machine section. Since the information about magnetic material is not provided by the manufacturer, it is assumed that magnetization characteristic for the stator and rotor cores is the 1020 Steel available from the software"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000108_j.biosystemseng.2019.08.002-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000108_j.biosystemseng.2019.08.002-Figure9-1.png",
        "caption": "Fig. 9 e Position of infrared distance measuring sensor connection points on the vehicle body at moment j.",
        "texts": [
            " The levelling system worked normally when the vehicle chassis travelled over continuously changing ground. The shorter levelling period, the lower the error but the greater the requirement for rapid hardware response and action. The initial time of each adjustment period was called moment j, and the time interval T aftermoment jwasmoment j\u00fe1(the time after moment j with interval T is moment j\u00fe1). The positions of the infrared distance measuring sensor connection points F1, F2, F3, and F4 on the vehicle body at moment j are as shown in Fig. 9. At the same time, the distance from connection points to point F1 in the vertical direction is \u25b3Hi (i \u00bc 1,2,3,4), which was calculated by the following equations: DH1 \u00bc 0 (1.a) DH2 \u00bc asinbc j (1.b) DH4 \u00bc bsinac j (1.c) DH3 \u00bcDH2 \u00fe DH4 \u00bc asinac j \u00fe bsinbc j (1.d) where, ac j is the pitch angle of the vehicle body at moment j; bc j is the roll angle of the vehicle body at moment j; a is wheel track of the vehicle chassis; b is the wheelbase of the vehicle chassis. If the suspension deformation is not considered, the distance from the wheel centre of the four wheels to point F1 at moment j\u00fe1 is Lij\u00fe1, it can be calculated by Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002781_j.matpr.2021.02.045-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002781_j.matpr.2021.02.045-Figure3-1.png",
        "caption": "Fig. 3. Force Applied.",
        "texts": [
            " In the form of a rectangle, the ends of the car bumper are formed. Table 1 Material properties of glass fiber and hybridized glass natural fiber reinforced composites. S.No Material Young\u2019s modulus GPa Poisson ratio Density Kg/m3 1 Glass fiber epoxy hybrid composites 45 0.25 2100 2 Hybridized glass fiber untreated Hemp fiber reinforced epoxy composites 62 0.28 1800 3 Hybridized glass fiber treated Hemp fiber reinforced epoxy composites 58 0.26 1600 The load is loaded into the front of the bumper. The Fig. 2 and Fig. 3 represents the boundary condition of the bumper. The meshing carried out here on the four wheeler bumper is a triangular mesh and the level of meshing is a fine mesh for the results will be as accurate as possible. The Fig. 4. represent the meshing of the bumper. Table 2 presents the Meshing details of glass fiber and hybridized glass natural fiber reinforced composite model bumper. Finite Element Analysis is the practical use of the Finite Element Method (FEM), which designers and researchers use to describe and solve complex primary, liquid, and multi-material science problem situations numerically and mathematically"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000084_012165-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000084_012165-Figure1-1.png",
        "caption": "Figure 1. Plain Cylindrical Hydrodynamic journal bearing with flexile liner",
        "texts": [
            " Series 1240 (2019) 012165 IOP Publishing doi:10.1088/1742-6596/1240/1/012165 There is a wide applicability of hydrodynamic plain cylindrical journal bearings. The stability analysis of such bearings is necessary from the designer\u2019s point of view. These bearings are lubricated with single oil film in which the wedge form of pressure distribution is generated. It applies forces between the mating surfaces of fluid film and bearing. The geometry of plain cylindrical journal bearing with flexible shell is explained in Fig. 1 below. The system is having a set of equations, which are solved to study and find the performance characteristics of journal bearing systems statically and dynamical point of view. According to Stokes micro-continuum theory The Reynolds equation for non-Newtonian fluids using Stokes micro-continuum theory in non-dimensional form is given by, (1) where, (2) (3) function G in the above expression is representing the interactions between base lubricant and additives. The effect of bearing flexibility in oil film thickness is given by, (4) where, (5) Term \u03b4 in equation (4) is the radial deformation component of bearing liner which is a function of deformation coefficient (Cd) and pressure, (6) where, (7) The non-dimensional form of a Reynolds equation for non-Newtonian fluid as described by using Stokes theory of micro-continuum can be expressed as (8) where, (9) (10) function G in the above expression is representing the interactions between base lubricant and additives"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000814_icarcv.2016.7838833-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000814_icarcv.2016.7838833-Figure1-1.png",
        "caption": "Fig. 1. Schematic of Non-collocated rotor and stator",
        "texts": [
            " Moreover, in contrast to continuous feedback controller that requires taking into account the nonholonomic constraints of the forklift in path generation stage, the MPMC is able to handle the nonholonomic constraints of the forklift in one optimisation loop. The computational burden due to the nonlinear MPC structure of path-following control is also reduced by successive online linearisation of the plant model and desired path function about the current operating point at each time step. Consider the kinematic model of a nonholonomic autonomous forklift with rear driving-steering wheel in the inertial frame x\u2212 y as (Fig. 1): \u23a1 \u23a2\u23a2\u23a3 x\u0307 y\u0307 \u03b8\u0307 \u03b1\u0307 \u23a4 \u23a5\u23a5\u23a6= \u23a1 \u23a2\u23a2\u23a3 cos\u03b8 cos\u03b1 0 sin\u03b8 cos\u03b1 0 1/R 0 0 1 \u23a4 \u23a5\u23a5\u23a6 [ v \u03c9 ] (1) where x and y are the forklift positions and \u03b8 is the forklift orientation in the Cartesian inertial frame, \u03b1 is the steering angle, v is the linear velocity of the driving wheel, \u03c9 is the the angular velocity of the steering wheel and R is the radius of the forklift instantaneous centre of rotation. R is also the radius of curvature, R= l/sin\u03b1 where l is the length of forklift. The distance between the two passive wheels at points B and C is defined as forklift track, d",
            " Consider the following linear timevarying discrete-time kinematic model of a nonholonomic forklift: \u03bek+1 = Ak\u03bek+Bkuk, (3) where \u03bek = [xk yk \u03b8k \u03b1k] T , uk = [vk wk] T and Ak,Bk are: Ak= \u23a1 \u23a2\u23a2\u23a3 1 0 \u2212Ts sin(\u03b8k)cos(\u03b1k)vk \u2212Ts cos(\u03b8k)sin(\u03b1k)vk 0 1 Ts cos(\u03b8k)cos(\u03b1k)vk \u2212Ts sin(\u03b8k)sin(\u03b1k)vk 0 0 1 Ts cos(\u03b1k)/lvk 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6 Bk = \u23a1 \u23a2\u23a2\u23a3 Ts cos(\u03b8k)cos(\u03b1k) 0 Ts sin(\u03b8k)sin(\u03b1k) 0 Ts sin(\u03b1k)/l 0 0 Ts \u23a4 \u23a5\u23a5\u23a6 , where k states the current time step and Ts is the sample time. The forklift system is subject to input and state constraints due to the limits for linear and angular velocity and accelerations of the driving wheel in addition to dynamic balance constraints as explained in the following section. The dynamic balance is considered in the form of the zero moment point (ZMP) [13] . The ZMP describes a point on a flat surface that no moments acting on the system. If the ZMP lies within the support polygon (SP), the convex polygon ABC in Fig. 1, then the system is dynamically balanced. The d\u2019Alembert formulation of the ZMP for the forklift shown in Fig. 1, assuming symmetric structure around y axis, can be represented as: xZMP = xCoG\u2212 zCoG( axb g )+ IGyz mg \u03c9z 2 (4) yZMP = yCoG\u2212 zCoG( ayb g )+ IGyz mg \u03b1z (5) zZMP = 0 (6) where (xCoG,yCoG,zCoG) is the position of centre of gravity (CoG) in body frame xb\u2212 yb\u2212 zb, IGyz is the second moment of inertia around yb\u2212 zb, axb and ayb accelerations in xb and yb direction, m is the overall mass of the forklift and load and \u03c9zb = \u03b8\u0307 and \u03b1zb = \u03b8\u0308 are the angular velocity and acceleration of the forklift around zb axis at point O"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003054_adem.202100557-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003054_adem.202100557-Figure1-1.png",
        "caption": "Figure 1. a) HEXRD setup for the in situ measurement of phase transformations during laser melting of a sample in the form of a thin platelet. The laser beammelts the upper part of the sample by moving along its upper edge and the occurring phase transformations are monitored simultaneously via X-ray diffraction. For this, the synchrotron X-ray beam is aligned perpendicular to both the sample platelet and the laser beam and directed on a stationary point in the upper center of the specimen. b) An investigated sample is displayed. The area, which has been laser-melted, is clearly visible on the upper edge. The thickness of the platelet is 400 \u03bcm. The used cross section of the impinging X-ray beam is indicated and marked in yellow. It has a height of 100 \u03bcm and a width of 400 \u03bcm.",
        "texts": [
            " Comparable cooling rates can be achieved and, thus, the chosen setup allows to investigate nucleation and growth mechanisms on an already solidified material of the same alloy. In summary, it should be noted that the used method makes it possible to measure solely inside the melt volume because the platelet is melted across its whole thickness. In the following, first results of in situ HEXRD experiments and corroborating microstructure investigations via scanning electron microscopy (SEM) prove the applicability of the modified setup for an accurate tracing of phase transformations upon solidification. The used experimental setup is schematically shown in Figure 1a. The laser and sample holder are contained in a chamber (not shown in this figure),[17] which is flooded with inert Ar gas. Thin platelets of the TiAl alloy with dimensions of 5 5mm and a thickness ranging from 200 to 500 \u03bcm can be mounted in the sample holder (Figure 1b). To melt the upper edge of the sample, a scanner unit moves the laser beam along a line on the edge of the sample (indicated with an arrow). The synchrotron X-ray beam, which is perpendicular to the platelet and the laser beam, has a height of 100 \u03bcm and a width of 400 \u03bcm and is directed on a stationary point just below the upper edge of the platelet. Therefore, phase transitions, which take place in an area up to 100 \u03bcm below the upper edge of the sample, can be detected. When the laser moves along the edge and melts the specimen, a full melting cycle (heating\u2013melting\u2013cooling) can be monitored and recorded",
            " To ensure that only liquid phase is detected during themelting of the platelet, i.e., a suitable melt pool size with respect to the X-ray beam is obtained, an extensive laser parameter study has Adv. Eng. Mater. 2021, 2100557 2100557 (2 of 7) \u00a9 2021 The Authors. Advanced Engineering Materials published by Wiley-VCH GmbH been conducted before the actual synchrotron experiments. The aim was to produce a melt pool that has a height larger than the height of the X-ray beam and that remains on the upper edge of the platelet without flowing down. Figure 1b shows the sample, which was investigated in this work. The picture was taken after the fusing experiment and the former melt pool is clearly visible on the upper edge of the specimen. A yellow area indicates the position and dimensions of the synchrotron X-ray beam, which was positioned solely in the melted area. For the experiment, an already additively manufactured Ti\u201348Al\u2013 2Cr\u20132Nb (at%) platelet-shaped specimen was chosen as base material to imitate the conditions of the real process as closely as possible",
            " The photon energy of the synchrotron X-ray beam equaled 100 keV and its cross section was 100 \u03bcm (height) 400 \u03bcm (width). For the experiment, a PILATUS 3X CdTe 2M area detector from Dectris AG, Switzerland, with a total number of pixels of 1475 1679 and a pixel size of 172 172mm2 was used. A detector frame rate of 50 Hz was chosen, which enables an exposure time of 0.02 s. The distance between the specimen and the detector was 1048mm. To investigate the phase transformations during laser melting via HEXRD, the platelet-shaped specimen was mounted on a specimen holder situated in a chamber (see Figure 1a), which is described in the study by Uhlmann et al.[17] The chamber was flooded with inert Ar gas. The measurement was performed in transmission geometry and the X-ray beam was positioned on the upper edge of the platelet, which was laser-melted during the experiment. From Figure 1a, it can be seen that the X-ray beam was directed on a stationary position perpendicular to the specimen as well as to the laser beam. For the laser melting Adv. Eng. Mater. 2021, 2100557 2100557 (6 of 7) \u00a9 2021 The Authors. Advanced Engineering Materials published by Wiley-VCH GmbH experiment, a laser power of 100W was selected, and the laser was moved along the upper edge of the specimen with a velocity of 20 mm s 1. The continuous wave ytterbium fiber laser YLR-400-AC from IPG Laser GmbH, Germany, with a wavelength of 1070 nm and a diameter of 60 \u03bcm was positioned at a working distance of 445mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001733_lra.2020.3003235-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001733_lra.2020.3003235-Figure3-1.png",
        "caption": "Fig. 3. Passive NIP Biped. Point mass at the hip, with pin joint at stance foot. Massless swing leg can be arbitrarily placed without dynamic implications.",
        "texts": [
            " 3) Ground impact is inelastic and the system does not rebound/bounce upon impact. 4) The stance foot does not slip and acts as an uncontrolled pin joint (some of the referenced papers use friction models and/or ankle control). The single support, or swing phase, of the legged NIP is given by the Lagrangian L L = T \u2212 U (1) T = ml2 2 \u03b8\u03072 U = mgl cos \u03b8, (2) where T and U are the system KE and PE, l is the length from the stance foot pivot to mass m, and Angle \u03b8 is the rotation of the pendulum per Fig. 3. The dynamic equations derived from (1) are shown to be d dt ( \u2202L \u2202\u03b8\u0307 ) \u2212 \u2202L \u2202\u03b8 = 0 (3) ml2\u03b8\u0308 = mgl sin \u03b8 (4) For state space x = [ \u03b8 \u03b8\u0307 ]T , the NIP double integrator can be rewritten as x\u0307 = [ \u03b8\u0307 g l sin \u03b8 ] (5) C. Impact Dynamics The swing leg impacts the ground with a rigid and nonconservative impact, wherein kinetic energy (KE) is lost from the system and an instantaneous change in pivot point occurs. Although energy is lost into the ground at foot contact, momentum is conserved about the foot touchdown point and thus the momentum about the swing toe before impact L\u2212 is equal to the angular momentum about the stance foot after impact L+, where the nomenclature \u2212,+ is used to describe information immediately before or after collision"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002541_0954407020974497-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002541_0954407020974497-Figure1-1.png",
        "caption": "Figure 1. Entire steering assembly.",
        "texts": [
            " The aspect of optimization is to achieve a steering geometry along with the wheel base and track width for a desired cornering performance with a wide range of turning radii, which in the case of a typical race car that hasn\u2019t been addressed yet. Moreover, these predicted optimal values ensure the achievement of fairly close result to the target and extend the possibility for further reduction of structural error. The typical rack and pinion steering system comprises a steering wheel, a steering column connected to the rack at one end. The rack is connected to a tie rod on each side as shown in Figure 1. The tie rods are pinned to the tie arms. The rotational input given to the steering wheel converts to translational motion of the rack through pinion. The translational motion of rack results in tyre rotation via the movement of tie-arm. The tyres are free to rotate about the King Pin Inclination (KPI) axis. For smooth turning of a vehicle, all four wheels should follow arcs having a common center, known as instantaneous center (IC). The Ackerman steering geometry intends to have such performance during cornering so as to avoid the tyres scrubbing or slip sideward"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000155_s12652-019-01522-9-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000155_s12652-019-01522-9-Figure1-1.png",
        "caption": "Fig. 1 Body-fixed frame and earth-fixed frame",
        "texts": [
            " A synchronous online optimal algorithm is proposed in this paper which can avoid the repetitive computation once the DP system changes and save a large amount of runtime which guarantees the realtime performance of the control scheme, and solves problems of energy conservation, emission reduction and equipment wastage decrease. This proposed DP optimal control law can maintain the vessel at desired states. The rest of the paper is organized as follows. The preliminaries are presented in Sect.\u00a02. In Sect.\u00a03, the time-based ADP method is employed for DP online optimal problem. Simulation results are given in Sect.\u00a04. Finally, the conclusion is given in Sect.\u00a05. The motion of DP vessel in three degrees of freedom can be modeled by the earth-fixed frame and the body-fixed frame 1 3 as shown in Fig.\u00a01 (Fossen 2011). OEXEYEZE is earth-fixed frame which can be seen as an inertial coordinate system. Original point OE is fixed on the surface of earth, axis XE points to the north. Axis YE points to the east and axis ZE points to the earths core. OBXBYBZB is body-fixed frame which is non-inertial. OB is fixed on the center of gravity, axis XB points to the prow. Axis YB points to the starboard and axis ZB points to the bottom. Both of two frames follow the right-hand rule. Since DP vessels always operate at a lower speed in the state of working generally, the low-frequency mathematical model of DP vessel can be shown as below (Fossen 2002b) In Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000876_0954407018824943-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000876_0954407018824943-Figure14-1.png",
        "caption": "Figure 14. Experimental setup: (a) belt transmission layout and (b) test bench schematic view and measurement principle.",
        "texts": [
            " Also, there are as many free belt-spans as there are wrapped belt-spans in a FEAD. For example, applying equations (40) and (41) to the FEAD in Figure 11, \u2018j\u2019 and naturally \u2018i\u2019 would range from 1 to 6. In addition, attention must also be paid to the first pulley during the analysis. In this case, the j 1 in equations (40) and (41) shall be replaced by np, where np is the number of pulleys composing the FEAD. Simulation and experimental results are presented for the serpentine belt drive system defined in Table 1 and presented in Figure 14(a), where the CS pulley is assimilated to the driving pulley, the two others AD1 and AD2 are accessory driven pulleys. The poly-V belt transmission test bench shown in Figure 14(a) and (b) is modular so that the position of the axes can be changed to reproduce almost any belt transmission layout. The driving axis and the driven axes are equipped with speed and torque metres (Magtrol, TM313, TM 310, TM 309) noted (MS) in Figure 14(b). The whole test bench is controlled via a control/data-acquisition system using National Instruments boards and Labview. The axis of the idler pulley is equipped with strain gauges to measure the belt tension in the slack span. The driving axis speed is controlled by a servomotor (Phase Automation), the resisting torques of the driven axes are also controlled by servomotors (Phase Automation) that serve as actuators to reproduce the loads applied by accessories in a FEAD. Given the sensitivities of the torque metres, the measurement error in terms of power is 60.93W at a speed (CS) of 600 r/min. The FEAD is considered to be running at 600 r/min and at 20 C. To analyse the evolution of PLFEAD (equation (1)) as a function of the FEAD power transmitted, a constant resisting torque TAD1 ranging from 5.5 to 45.3Nm is applied on the shaft AD1 (Figure 15). Consequently, as the CS in Figure 14, speed is controlled to maintain the speed at 600 r/min, to offset TAD1 a driving TCS and a resisting TAD2 torques from 1.5 to 2.1 TAD1 and 0.5Nm are measured on the shafts CS and AD2 (Table 1), respectively. Next, two case studies, S1 (low torque) and S2 (high torque) are considered in Figure 15. Changing from situation S1 to S2, the objectives are (1) to validate globally the power loss models in equation (1) and (2) to highlight the effectiveness of the PLslip modelling, since except PLslip, the other models in equation (1) have already been either analysed, for example, PLhys, 10,11 or assumed, by default, to be validated, for example, PLbear from SKF",
            "13 The belt used here has the same characteristics as that in Silva et al. except its length (Table 2). Moreover, it is installed with a setting-tension T0 =204N which is far below a common setting-tension representative of Truck FEAD applications (T0 =600N) (Figure 1). However, as previously mentioned, a lower settingtension permits highlighting new results (PLslip) not verified yet in the previous analyses.10,11 The FEAD power losses represented by equation (1) were both measured and simulated, with the model presented, see Table 3. Since the test bench in Figure 14 is equipped with a speed and torque measuring system on each shaft, the power supplied by the CS and the power consumed by the accessory drives are known. The difference between the power supplied and consumed is the power lost/measured in the FEAD, PLM in Table 3. In Table 3, the simulated results are obtained from the power loss models of equation (1) with the exception of PLvib and PLtens hys for which only the power loss models are given here. In case of need, material properties from DMA characterization of the beltelastomer and the belt-cords can be found in Silva et al",
            " On the contrary, PLhys remains almost unchanged as the belt-hysteresis depends mostly on the belt-bending10,11 and indirectly on the FEAD geometry which is identical in both cases S1 and S2. From S1 to S2, the bearings of the shafts CS and AD1 are more loaded and those of AD2 and IDLER less loaded, it was verified numerically and experimentally. Hence, for AD2 and IDLER, bearing power losses decrease compared to S1 (S2, Figure 17). Moreover, from S1 to S2, the slip power losses have been increasing (and highlighted) on AD1 and CS providing a new power loss distribution of the FEAD considered in Figure 14. In Figure 15, for both experimental and simulation results, three zones can be identified: (A) where the setting-tension T0 is enough to transmit power on all shafts. In this case, the FEAD total power losses Table 2. Poly-V belt geometric parameters and data. Length: 1715 mm Ht : 0.9 mm Density: 0.16 kg/m Hb : 1.6 mm Angle a: 40 Hc : 2.4 mm Ribs n: 10 Bd : 3.56 mm EPDM: ethylene propylene diene monomer; PET: poly ethylene terephthalate. Table 4. Distribution of the simulated power losses in Table 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure17-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure17-1.png",
        "caption": "Fig. 17. Virtual prototype model of 3-R(RRR)R+R HAM: (a) Virtual prototype model; (b) explosion diagram.",
        "texts": [
            " The workspace of the 3-R(RRR)R+R HAM in the cartesian coordinate system can be obtained, as shown in Fig. 16. According to Fig. 16 (a), (b) and (c), when the azimuth range of 3-R(RRR)R+R HAM is 0\u2013360\u25e6 continuous rotation and the pitch angle is 0\u2013390\u25e6 continuous rotation, the 3-R(RRR)R+R HAM can achieve the satellite tracking under the antenna pitch and azimuth motion. A prototype of 3-R(RRR)R+R HAM is made with a ratio of 1:2 to verify the mobility capability of the HAM. The virtual prototype model and explosion diagram of 3-R(RRR)R+R HAM are shown in Fig. 17. From the virtual prototype model and explosion diagram, the structure and assembly relationship can be shown clearly. Based on the virtual prototype model of the 3-R(RRR)R+R HAM, the parts of the HAM are processed and assembled. The physical prototype of the 3-R(RRR)R+R HAM is shown in Fig. 18. The motion experiment of 3-R(RRR)R+R HAM prototype is carried out to verify the motion performance of 3-R(RRR)R+R HAM. To facilitate the observation and removal of the antenna reflector connected to the polarization rotation mechanism, for the 3-R(RRR)R+R HAM prototype, 90\u25e6 pitch motion and 45\u25e6 pitch angle 360\u25e6 azimuth motion are, respectively, carried out and the motion experimental results are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003884_20.717751-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003884_20.717751-Figure6-1.png",
        "caption": "Fig 6 The mesh of the 3D skewed slots structure",
        "texts": [
            " 4 At no load, the normal component of magnetic flux density For the nominal speed (1450 r.p.m), we can see in figure 5 the steady state torque waveform obtained by 2D and 3D approaches. In this figure, we can see also that both approaches give the same results. From this study, we can conclude that the 3D model is validate by a comparison with 2D model for the same computation conditions. B. Skewed Case For the skewed rotor bars with 3D calculation, it is necessary to model a quarter of machine. In this case, the number of unknowns is 190127. We can see in figure 6 the studied mesh. It may be noted that the short circuit rings are included in magnetodynamic model. For 2D approach, we use 9 disks with a total of5898 unknowns. The machine has been modelled at locked rotor. We can compare in figure 7 the 2D and 3D models, for the transient armature current (Fig. 7.a.). The corresponding electromagnetic torque calculation is given on figure 7.b. For this simulation, the numerical results are in good agreement. We can note that the calculation realised with 9 discs gives a satisfactory results compared with 3D approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001690_s12046-019-1263-1-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001690_s12046-019-1263-1-Figure4-1.png",
        "caption": "Figure 4. Locations of accelerometers: (a) radial direction and (b) axial direction.",
        "texts": [
            " The accelerometers have been mounted permanently, perpendicular to each other on the input shaft bearing housings to minimize transmission path effects as Table 1. Equation of statistical parameters. Mean x \u00bc 1 N PN i\u00bc1 xi Median Median RMS RMS \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 N PN i\u00bc1 \u00f0x2 i \u00de s Standard deviation r \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 N PN i\u00bc1 \u00f0xi x\u00de2 s Peak to peak PP \u00bc max min Crest factor FC \u00bc PP RMS Skewness Sk \u00bc 1 N PN i\u00bc1 \u00f0xi x\u00de3 r3 Kurtosis Kr \u00bc 1 N PN i\u00bc1 \u00f0xi x\u00de4 r4 seen in figure 4. A PCB-480C02-type signal conditioner is used to strengthen the accelerometer output, as shown in figure 5. The vibration signals, which come from the accelerometer on the worm gearbox, are transferred to a computer via NI (which is a trademark of National Instrument) DAQ Card 6036E. Also, this analyser acquired pulse signals of the inductive sensor, which produces one pulse for each rotation of the motor shaft. The vibration and angular position signals acquired from accelerometers and inductive sensor are sampled at an appropriate rate and recorded on the computer using a NI data acquisition system and LabVIEW 7",
            " Statistical features are shown in figures 8 and 9 for two perpendicular sensors. As seen in the figures, the circle marks represent the data from the radial direction sensor and triangle points represent the data from the axial direction. As can be seen in figures 8 and 9, most of the statistical features in time domain do not display a reliable trend with the progression of the fault, but some of them show an increasing trend for the averaged worm gear vibration in the radial direction, as shown in figure 4(a). RMS, peak to peak and standard deviation values show steady increase, with respect to progression of the failure. On the other hand, all statistical properties of averaged vibration signals in time domain, which are acquired from axial direction as shown in figure 4(b), show inconsistent trend with the progression of the fault. Figure 10(a) illustrates the spectrum representation of averaged worm gear vibration in the radial direction. The rotating frequency of worm gear input shaft can be seen as 1X in this figure. This is seen even in healthy worm gear but it is not the unusual situation for worm gear according to former studies [2, 11]. The amplitude of the rotating frequency of worm gear gradually increases with the progression of fault. Also, the amplitude of frequency increases steadily in the higher frequency range, which is caused by increased friction [2]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001293_s00170-019-04519-y-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001293_s00170-019-04519-y-Figure2-1.png",
        "caption": "Fig. 2 Thread tapping device",
        "texts": [
            " This latter is characterized by a section with four lobes, an active truncated part with an angle of 10\u25e6 over 3 times the pitch and a maximum diameter of 8.21 mm. According to the manufacturer, this overabundance of dimension compared with the nominal diameter is to compensate the elastic return after the thread forming and the disengagement of the tap. The geometry and the tolerances of cut and form taps were accurately detailed in the works of Landeta et al. [15]. To ensure a perpendicular position between the thread helix axis and the washer plane, a specific device was used to maintain the tap in a perpendicular position (Fig. 2). Initially, a centering pin was used to place properly the washer on the adopted device. After clamping firmly the washer, a guide component with a helical groove was also considered to guide the fluteless cutting tap or the forming tap (Fig. 2). Both guide components were machined for each tapping process with a strict geometrical conditions. The only difference between these guide components is at their helical groove. For the cut tapping process, the helical groove is limited for the passage of the fluteless tap. For the Fig. 1 a Squares. b Washers. c Annealed washers form tapping process, it has the same shape with the thread but with a very low height. The geometry was studied through the analysis of a radial section before and after the thread realization"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000492_s12206-016-0115-8-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000492_s12206-016-0115-8-Figure3-1.png",
        "caption": "Fig. 3. Line of action of the sun gear ~ planet gear mesh.",
        "texts": [],
        "surrounding_texts": [
            "motion of each element at the center points and the bearing force between the planet gear and the carrier.\n1 2 1 a a ad d b \u03b8= + \u00b4 uur uur uurr , 0 1 1 a a a\nbd d \u03b8 2 = + \u00b4\nruur uur uur (3)\n0 1 2 a a a\u03b8 \u03b8 \u03b8 .= = uur uur uur\n(4)\nIn this study, the equation of motion for each component is organized in a vector format. Mesh forces ( spF , prF ) are transferred along the unit vector in the direction of the line of action. The unit vector in the direction of the line of action was aligned along the directions from the sun gear to the planet gear, and from the planet gear to the ring gear. In the carrier and planet gear relationship, the bearing force is transferred along the horizontal, vertical unit vector ( cpn ) of piC . In addition, the vibration model of a tooth contact point defined the equivalent modulus of elasticity of the tooth contact point when the tooth profile curve of the center surface of a gear passes the 1 pitch point with consideration of the elastic deformation of a tooth [15]. The load distributed to and transferred along the tooth width was defined as the concentrated force applied to a pitch point of a crowned tooth surface, and the friction force between the tooth surfaces was neglected. The equations of motion of each component are detailed below.\n2.1 Sun gear\nFig. 2(a) illustrates the motion relationship between a sun gear and a planet gear. Eqs. (5) and (6) are the equations of motion of stations \u20181\u2019 and \u20182\u2019 of the sun gear, respectively. The mesh force spF uur from the engagement of the sun gear\nand the planet gear is applied to the sun gear. The translational\ndisplacement ( 0 1 1 ss s \u03b8\u00b4\nuuur uur ) and moment ( 2 1 1 ss s F\u00b4 uuur uur ) are addition-\nally generated from the rotational displacement ( 1 s\u03b8 uur ). In addition, the location ( 2s a uuur\n) of the mesh force applied on the sun gear body is identified so that the moment generated from the mesh force can be considered. Each vector component is shown in the Appendix.\n(5) . (6)\n2.2 Planet gear\nFigs. 2(a)-(c) illustrate the motion relationships of the planet gear and the other elements. Eqs. (7) and (8) are the equations of motion for stations \u20181\u2019 and \u20182\u2019 of the planet gear. The carrier is simultaneously engaged with the sun gear and the ring gear, and is coupled with the planet gear through bearings. Therefore, the mesh forces ( spF uur , prF uur ), bearing force ( cpF uur\n), and moment ( cpM uuuur ) from the coupling of the sun gear and the\nring gear are applied to the planet gear. The mesh force with the sun gear is tangential at body \u2018b\u2019 of the planet gear, and the mesh force with the ring gear is tangential at \u2018c\u2019. The force and moment from the bearing on the carrier are applied at the center point ( 0 0p ,c ) of each body.\n(7)\n(8)\n2.3 Ring gear\nFig. 2(b) illustrates the motion relationship between the ring gear and the planet gear. As the ring gear is engaged with the planet gear, the mesh force prF uur is applied on point \u2018d\u2019 on the\nbase circle. The additional displacement ( 0 1 r 1r r \u03b8\u00b4\nuuur uur ) and mo-\nment ( 2 1 1 rr r F\u00b4\nuuur uur ) are generated from the equation of motion\ndue to the influence of the tooth width of the ring gear as well. Eqs. (9) and (10) are the equations of motions between stations \u20181\u2019 and \u20182\u2019 of the ring gear.\n(9) . (10)\n2.4 Carrier\nFig. 2(c) illustrates the motion relationship between the planetary gear and the carrier. The carrier is coupled with the planet gear through bearings. Bearing force ( cpF uur ) is\napplied on the carrier, and an additional moment ( 2 1 1 cc c F\u00b4\nuuur uur )\nis generated at station '2' due to the width of the carrier and the force from station '1'. The bearing force ( cpF uur ) is applied\non the center point ( 0 0p ,c ) of the carrier and the planetary gear. The corresponding relationships are shown in Eqs. (11) and (12).\n(11) . (12)\n2.5 Mesh force\nThe mesh forces ( spF uur , prF uur ) between the sun gear and the\nplanet gear and between the planet gear and the ring gear are transferred through the unit vectors along the lines of action ( spn uuur , prn uuur\n), and their magnitudes ( spF , prF ) are determined by the tooth displacements ( sp\u03b4 , pr\u03b4 ) and tooth stiffness ( spk , prk ). The tooth stiffness values are the equivalent stiffnesses of the gear teeth. spe , pre are the transmission errors\n.",
            "on the lines of action. \u00b6spe ,\u00b6pre represent the magnitudes of\nthe transmission errors. In Eqs. (15) and (16), l represents the order of the harmonic component, \u03c9 represents the tooth passing frequency, sp\u03a6 , pr\u03a6 represent the phase difference\nbetween a planet gear in an arbitrary position and the reference planet gear, and rs\u03a6 represents the phase differences between the reference sun gear ~ planet gear mesh and the planet gear ~ ring gear mesh [16]. The magnitude of the transmission error is a source of vibration on the gear teeth surface, which acts as a source of vibration in a forcedvibration analysis.\n( )sp sp sp sp sp sp spF F n k \u03b4 e n= = -\nuur uuur uuur (13) ( )pr pr pr pr pr pr prF F n k \u03b4 e n= = - uur uuur uuur (14)\n\u00b6 ( )spl \u03c9t \u03a6 spl spe e e + = (15)\n\u00b6 ( )pr rsl \u03c9t \u03a6 \u03a6 prl pre e e . + + = (16)\nspn uuur , prn uuur\nare expressed as Eqs. (17) and (18), due to points \u2018a\u2019, \u2018b\u2019, \u2018c\u2019, and \u2018d\u2019, which are the tangential points of the lines of action and the base circle in Figs. 3 and 4. The line of action becomes tangential with the base circle of each gear as the line of action rotates as much as the base circle helix angle about the pitch point on the action plane. The position vector of each point can be described by the base circle radius ( cr ), the operation pressure angle ( sp pr,f f ), the tooth width (b), the base circle helix angle (\u0393) , the contact center distance ( meshC ), and the planet gear position angle (\u03c8) . Eqs. (19)- (26) show the position vectors of \u2018a\u2019, \u2018b\u2019, \u2018c\u2019, and \u2018d\u2019.\nsp abn ab =\nuuruuur uur (17)\npr cdn cd =\nuuruuur (18)\n1 1 1 1ab as s b s a s b= + = - + uur uur uur uuruur = 1 1 1 1s a s p p b- + + uur uuur uuur\n(19)\n( ) ( ) ( ) ( )1 s sp s sp s sp bs a r cos \u03b1 i r sin \u03b1 j r tan tan \u0393 k 2 \u00e6 \u00f6= + + + f\u00e7 \u00f7 \u00e8 \u00f8\nuur r r r\n(20)\n( ) ( )1 1 mesh meshs p C cos \u03c8 i C sin \u03c8 j= +\nuuur r r (21)\n( ) ( ) ( ) ( )1 p sp p sp p sp bp b r cos \u03b1 i r sin \u03b1 j r tan tan \u0393 k 2 \u00e6 \u00f6= - - + - f\u00e7 \u00f7 \u00e8 \u00f8\nuuur r r r\n(22) 1 1 1 1 1 1 1 1cd cp p d p c p d p c p r r d= + = - + = - + + uuur uuur uur uuur uur uuur uuruur (23)\n( ) ( ) ( ) ( )1 p pr p pr p pr bp c r sin \u03b1 i r cos \u03b1 j r tan tan \u0393 k 2 \u00e6 \u00f6= + + - f\u00e7 \u00f7 \u00e8 \u00f8\nuur r r r\n(24)\n( ) ( )1 1 mesh meshp r C cos \u03c8 i C sin \u03c8 j= - -\nuuur r r (25)\n( ) ( ) ( ) ( )1 r pr r pr r pr br d r sin \u03b1 i r cos \u03b1 j r tan tan \u0393 k 2 \u00e6 \u00f6= + + - f\u00e7 \u00f7 \u00e8 \u00f8\nuur r r r .\n(26)\nTooth displacement on a line of action is determined by the body displacement at each gear element, \u2018a\u2019, \u2018b\u2019, \u2018c\u2019 and \u2018d\u2019. The tooth displacement between the sun gear and the planet gear ( sp\u03b4 ) is determined by the sun gear displacement ( s\u03be ) at point \u2018a\u2019 and the planetary gear displacement ( ps\u03be ) at point \u2018b\u2019. The tooth displacement between the planetary gear and the ring gear ( pr\u03b4 ) is determined by the planetary gear displacement ( pr\u03be ) at point \u2018 c \u2019 and the ring gear displacement ( r\u03be ) at point \u2018d\u2019. Eqs. (27)-(32) express the tooth displacement of each gear element.\nsp s ps\u03b4 \u03be \u03be= - (27)\npr pr r\u03b4 \u03be \u03be= + (28)\n1 1 s s s sp\u03be d \u03b8 s a n\u00e6 \u00f6= + \u00b4 \u00d7\u00e7 \u00f7\n\u00e8 \u00f8\nuur uuruur uuur (29)\n1 1 ps p p sp\u03be d \u03b8 p b n\u00e6 \u00f6= + \u00b4 \u00d7\u00e7 \u00f7\n\u00e8 \u00f8\nuur uuuruur uuur (30)\n1 1 pr p p pr\u03be d \u03b8 p c n\u00e6 \u00f6= + \u00b4 \u00d7\u00e7 \u00f7\n\u00e8 \u00f8\nuur uuruur uuur (31)\n1 1 r r r pr\u03be d \u03b8 r d n .\u00e6 \u00f6= + \u00b4 \u00d7\u00e7 \u00f7\n\u00e8 \u00f8\nuur uuruur uuur (32)",
            "2.6 Bearing force and moment\nThe support bearing of a planet gear enables the bearing displacement to be predicted from the displacements of the planet gear and the carrier. Since the carrier and the planet gear are not coaxial, the translational displacement ( cp\u03b4 uuur ) of the bearing is computed from the translational ranges of the bearing and carrier, the rotational range of the carrier, and Eqs. (33) and (34). The bearing force due to the bearing stiffness and bearing translational displacement is applied along the unit vector ( cpn uuur ), and its magnitude ( cpF ) is computed from Eq. (35). In addition, for the support bearing of the planetary gear, a moment is generated as shown in Eqs. (37)-(40) due to the rotational displacement ( cp2\u03b4 ) of the planetary gear. Here, rk , ak and tk are the radial stiffness, axial stiffness, and tilting stiffness of the planet gear support bearing, respectively.\n1 1 1 0 1 cp c c p cp,x cp,y cp,z\u03b5 d \u03b8 c p d \u03b5 \u03b5 \u03b5= + \u00b4 - = + + uur uur uuuur uuruur uuuur uuuur uuuur\n(33)\n( )cp r cp,x cp,y a cp,z\u03b4 k \u03b5 \u03b5 k \u03b5= + + uuur uuuur uuuur uuuur\n(34)\ncp cp cpF \u03b4 n= \u00d7 uuur uuur\n(35)\ncp cp cpF F n= uur uuur\n(36)\n1 1 cp c p cp2,x cp2,y cp2,z\u03b5 \u03b8 \u03b8 \u03b5 \u03b5 \u03b5= - = + + uur uuruur uuuuur uuuuur uuuuur\n(37)\n( )cp2 t cp2,x cp2,y\u03b4 k \u03b5 \u03b5= + uuuur uuuuur uuuuur\n(38)\ncp cp2 cpM \u03b4 n= \u00d7 uuuur uuur\n(39)\n1 cp cp pM M \u03b8 .=\nuuruuuur (40)\n2.7 Planetary gear system local transfer matrix\nRearranging the equations of motion provided above yields the local transfer matrix of the planetary gear system. When the number of planet gears increases, the additional planetary gears can be considered from Eqs. (3)-(12), and can be derived using the previously employed method. The equation of motion in vector notation can be rearranged to linear algebra notation according to Table 1.\ng[T ] represents the transfer matrix of the planetary gear system, and 0g[T ] , 1g [T ] , and 2g [T ] represent the stiffness, damping, and inertial elements, respectively. g[F ] is a vari-\nable that represents the transmission error excitation force. S\u00e9 \u00f9\u00eb \u00fb is a variable that represents the translational and rotational\ndisplacements, force, and moment of each element. Variables in bold represent column variable matrices. Details of the transfer matrix elements of the planetary gear expressed in Eqs. (41)-(43) are explained in the appendix.\nn n n n n[S ] [d ,\u03b8 ,F ,M ]= (41) n 1 n n n\ng g[S ] [T ][S ] [F ]+ = + (42) 2 g 0g 1g 2g[T ] [T ] [T ]\u03bb [T \u03bb .]= + + (43)\n3. Solution technique of the transfer matrix method\n3.1 Local transfer matrix derived by the Hibner branch method\nModeling using HBM for the local transfer matrix is necessary to derive the global transfer matrix. For multi-axial systems such as external and planetary gears, the derivation must consider the coupling between each axis, but there are cases where empty spaces exist, as shown in Fig. 5. There are a total of 2 lines, 3 elements, and 4 stations. Element#3 of line#1 and element#1 of line#2 are empty spaces with no axis. Analysis of station#3 of element#2, line#1 revealed that it is equivalent to stations#3 and #4 of element#3, and the analysis of stations#1 and #2 of element#1, line#2 showed that it is equivalent to station#2 of element#2. When the state variables of lines#1 and #2 of station#1 are 1[ 1 ]S and 1[ 2 ]S , respectively,\nthe total state variable of station#1 is 1[ ] =S 1 1[[ 1 ],[ 2 ],1]TS S . Line#1 of element number# 1 in Fig. 5 is empty, but when the state variable of station#2 of line#2 is made equivalent up to station#1, local transfer matrix modeling of element number 1 is possible even when there is empty space. Because of this, modeling using the transfer matrix method is possible for complex, multi-axial gear systems.\n2 1 1\u00e9 \u00f9 \u00e9 \u00f9 \u00e9 \u00f9=\u00eb \u00fb \u00eb \u00fb \u00eb \u00fbS T S (44)\n2 1\n2 1 1 10 0 2 0 0 2 . 1 10 0 1 \u00e9 \u00f9 \u00e9 \u00f9\u00e9 \u00f9\u00e9 \u00f9 \u00e9 \u00f9 \u00e9 \u00f9\u00eb \u00fb \u00eb \u00fb \u00eb \u00fb\u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa=> = \u00e9 \u00f9 \u00e9 \u00f9 \u00e9 \u00f9\u00eb \u00fb \u00eb \u00fb \u00eb \u00fb\u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa\u00e9 \u00f9 \u00e9 \u00f9\u00eb \u00fb \u00eb \u00fb\u00eb \u00fb\u00eb \u00fb \u00eb \u00fb T T S ST S I S"
        ]
    },
    {
        "image_filename": "designv11_14_0000569_s11029-016-9578-z-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000569_s11029-016-9578-z-Figure1-1.png",
        "caption": "Fig. 1. Geometry of a rectangular laminated composite plate.",
        "texts": [
            " The objective of this paper is to employ Reddy\u2019 third-order shear deformation theory to analyze the bending of a laminated plate by using a simple isoparametric four-node finite element with five degrees of freedom at each node. 2. Kinematics According to Reddy\u2019s third-order shear deformation theory (TSDT) [17], the displacement field can be expressed as u u z z h w xx x1 34 3 = + \u2212 + \u2202 \u2202      \u03c8 \u03c8 , u v z z h w yy y2 34 3 = + \u2212 + \u2202 \u2202      \u03c8 \u03c8 , u w3 = , where u v w x y, , , ,\u03c8 \u03c8and are five unknown midplane displacement functions of the plate, and h is its thickness (see Fig. 1). The linear strains associated with the displacement field are \u03b5 \u03b5 \u03b5 \u03ba \u03ba1 1 1 2 1 2= = + +xx z z ( ), \u03b5 \u03b5 \u03b5 \u03ba \u03ba2 2 2 2 2 2= = + +yy z z ( ), \u03b5 \u03b53 0= =zz , \u03b5 \u03b5 \u03b5 \u03ba4 4 2 4 2= = +yz z , (1) \u03b5 \u03b5 \u03b5 \u03ba5 5 2 5 2= = +xz z , \u03b5 \u03b5 \u03b5 \u03ba \u03ba6 6 6 2 6 2= = + +xy z z ( ), where \u03b5 \u03ba \u03c8 \u03ba \u03c8 1 1 1 2 2 2 2 4 3  = \u2202 \u2202 = \u2202 \u2202 = \u2212 \u2202 \u2202 + \u2202 \u2202       u x x h x w x x x, , , \u03b5 \u03ba \u03c8 \u03ba \u03c8 2 2 2 2 2 2 2 4 3  = \u2202 \u2202 = \u2202 \u2202 = \u2212 \u2202 \u2202 + \u2202 \u2202       v y y h y w y y y, , , \u03b5 \u03c8 \u03ba \u03c84 4 2 2 4  = + \u2202 \u2202 = \u2212 + \u2202 \u2202      y y w y h w y , , \u03b5 \u03c8 \u03ba \u03c85 5 2 2 4  = + \u2202 \u2202 = \u2212 + \u2202 \u2202      x x w x h w x , , \u03b5 \u03ba \u03c8 \u03c8 \u03ba \u03c8 \u03c8 6 6 6 2 2 24 3 2 = \u2202 \u2202 + \u2202 \u2202 = \u2202 \u2202 + \u2202 \u2202 = \u2212 \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 \u2202 u y v x y x h y x w x x y x y, , y       "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002533_j.tws.2020.107247-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002533_j.tws.2020.107247-Figure4-1.png",
        "caption": "Fig. 4. The selected finite element model of pyramid lattice structure with hollow truss.",
        "texts": [
            "11 platform) was carried out to analyze the effect of temperature on compression behavior compressive behavior and the corresponding compressive failure mechanisms of the hollow lattice truss sandwich structures. It suffices to consider only a single tube of the pyramidal unit cell. The top and bottom face sheets were both taken as rigid. It is assumed that the pyramidal core is perfectly bonded to the rigid faces and the end faces of each inclined tube approach each other by the compressive displacement \u03b4 in the 3 direction in Fig. 4. All degrees of freedom of the bottom rigid face sheets are fixed. The 3-direction displacement was unconstrained on the moving top face sheet while rotations about all axes were fixed resulting in a macroscopic state of compression on the structures being analyzed. The geometry was discretized using primarily linear, 4- node, selective reduced integration shell elements (S4R). The mesh convergence test was carried out. The mesh grid was one over ten of the diameter of hollow truss. No imperfections were introduced to the simulated structures, whereas the structures in the physical experiments possessed a number of defects that tended to reduce their compressive modulus and strength"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002110_nap51477.2020.9309696-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002110_nap51477.2020.9309696-Figure3-1.png",
        "caption": "Fig. 3. Residual equivalent stresses in the studied models.",
        "texts": [
            " The smallest values of equivalent stresses of about 9 MPa occur at the areas of sharp change in the cladding trajectory, which are typical for hollow triangle and prism models. This result is attributable to the geometric shape of the samples (the presence of angles) and the redistribution of stresses during cooling. The level of stresses in the residual state in the prototypes after cladding of three layers corresponds to 400-440 MPa, that does not exceed the strength limits of the substrate material and welding wire. Fig. 3 shows the fields of distribution of equivalent stresses for different models (samples). The stress level in the substrate varies in the range of 120-350 MPa and reaches the maximum value in the area of fusion of the substrate and the deposited metal, which is a typical phenomenon for all simulated samples. Moreover, the distribution of residual stresses in the vertical printed prototypes is uneven: the maximum stress level of about 500 MPa is typical for the first deposited layer, and the minimum of 390 MPa for the last layer, that is attributed to less thermal impact on the layer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000033_physrevfluids.4.063304-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000033_physrevfluids.4.063304-Figure1-1.png",
        "caption": "FIG. 1. Experimental setup. (a) The suspension sample is placed in a cylindrical container. An impactor motivated by an actuator is placed above the suspension. An ultrasound transducer is placed under the container, held by an optical stage. (b) Examples of the B-mode images obtained at different y positions by moving the ultrasound transducer. Relevant dimensions of the setup are labeled in the figure.",
        "texts": [
            " We apply this method to image the transient flows inside dense suspensions that result from impact at different incident angles. This allows us to address the question to what extent these flows, and the associated jamming fronts, retain axisymmetry along the propagating direction and how they deform when approaching a solid boundary. Our experiment studies impact at a suspension-air interface through ultrasound imaging and reconstruction of the resulting flow in three dimensions. The experimental setup is schematically illustrated in Fig. 1. The suspension was contained in a cylindrical vessel with an inner diameter of 10 cm. The impactor was driven by a linear actuator (SCN5, Dyadic Systems) mounted above the container. The incident angle \u03b8I,1 defined as the angle from the \u2212z axis to the incident direction, and the impact speed Up were both adjustable. The plane in which the impactor rotated and moved was aligned with the center of the container, and we define this plane as the y = 0 plane. The head of the impactor was a hemisphere, so it contacted the suspension surface in the same manner regardless of \u03b8I",
            " 063304-2 the impactor reached a position 5 mm above the suspension surface (the Verasonics ultrasound system we used is capable of imaging up to 10 000 frames per second). Each acquisition generated a 2D slice in the x-z plane. For vertical impact, 2D images at y = 0 mm (directly below the impactor) are sufficient to reconstruct the 3D flow field because of its rotational symmetry [24]. In order to visualize nonaxisymmetric 3D flow under tilted impact, we moved the transducer along the y axis as shown in Fig. 1(b) and combined the 2D slices obtained from different y positions. At each \u03b8I, the transducer was moved from y = 0 mm to y = 20 mm, in increments of 5 mm. At each y, the impact experiment was repeated three to nine times. After every impact, the suspension was fully relaxed by gently shaking and rotating the container. The accumulated data were used to reconstruct an averaged 3D flow field for each \u03b8I. III. IMPACT-ACTIVATED FRONTS To obtain the flow field, we used reconstructed B-mode images with trackable speckle patterns, whose motions represent the flow"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003670_60.790892-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003670_60.790892-Figure5-1.png",
        "caption": "Fig. 5. Phaser diagram for unity power factor operation",
        "texts": [
            "00 0 1998 IEEE The above analysis shows that unity pf can be obtained at two different values of Xc2 for a particular value of slip. When the circuit is operated at one of these two values of X,,, the fictitious capacitor will have some voltage Vs2 across it with a phase angle 8 with respect to the main supply voltage V, ,. If the PWM voltage inverter can be controlled to produce the same output voltage V,, 427 terminals. Hence, the effect of PWM inverter can be represented by a variable capacitor Cz in the equivalent circuit as shown in Fig. 3. A phasor diagram of the proposed scheme for unity pfis shown in Fig. 5. The hnction of the PWM inverter is to produce a voltage Vs2 with an appropriate phase and angle 0 so that the current through the auxiliary stator winding Is2 has a leading reactive component Is2 sin 0 of the phasor current I, 12 drawn by the motor in the absence of the auxiliary stator winding. The reactance Xc2 is the ratio of the PWM output voltage V,, and the current through the auxiliary winding Isz, where Vs2 is the fundamental component of the PWM output voltage. In practice, the PWM inverter output voltage has a kndamental as well as harmonic components",
            " Because of the narrow range over which the efficiency of the machine varies with power factor (Fig. 7(b)), the machine will still operate close to the maximum efficiency. CONTROL STRATEGY The PWM voltage inverter has to produce a particular value of VS2 (1) with a particular value of phase angle 0 in order to maintain the desired p$ Hence, at a particular load, the required magnitude and phase of Vs2 (1) has to be obtained and kept constant with the help of a controller. Referring to the phasor diagram, Fig. 5 , active power associated with the auxiliary stator winding is given by: PS2 = E S 2 Is2 sin (0 - P) Wlphase (9) The auxiliary winding receives this power from the main stator winding by transformer action. Current through the auxiliary winding is given by Control strategy can be realized with the help of eqns. (9) and (10). Reactive power in the main stator winding can be sensed and compared with the desired reactive power, Qref, and the error signal thus obtained can be used to operate the controller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000340_978-94-017-7300-3_10-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000340_978-94-017-7300-3_10-Figure2-1.png",
        "caption": "Fig. 2 Left: Typical morphologies of ribbons subjected to twist and stretching: (a) helicoid, (b, c) longitudinally wrinkled helicoid, (d) creased helicoid, (e) formation of loops and self-contact zones, (f) cylindrical wrapping, (g) transverse buckling and (h) twisted towel shows transverse buckling/wrinkling. Right: (i) Experimental phase diagram in the tension-twist plane, adapted from [3]. The descriptive words are from the original diagram [3]. Note that the twist used in the experiment is not very small; this apparent contradiction with our hypothesis \u03b7 1 (Eq. (2)) is clarified in Appendix A",
        "texts": [
            "1 Overview A ribbon is a thin, long solid sheet, whose thickness and length, normalized by the width, satisfy: thickness: t 1; length: L 1. (1) The large contrast between thickness, width, and length, distinguishes ribbons from other types of thin objects, such as rods (t \u223c 1, L 1) and plates (t 1, L \u223c 1), and underlies their complex response to simple mechanical loads. The unique nature of the mechanics of elastic ribbons is demonstrated by subjecting them to elementary loads\u2014twisting and stretching\u2014as shown in Fig. 1. This basic loading, which leads to surprisingly rich plethora of patterns, a few of which are shown in Fig. 2, is characterized by two small dimensionless parameters: twist: \u03b7 1; tension: T 1, (2) where \u03b7 is the average twist (per length), and T is the tension, normalized by the stretching modulus.1 Most theoretical approaches to this problem consider the behavior of a real ribbon through the asymptotic \u201cribbon limit\u201d, of an ideal ribbon with infinitesimal thickness and infinite length: t \u2192 0,L \u2192 \u221e. A first approach, introduced by Green [1, 2], assumes that the ribbon shape is close to a helicoid (Fig. 2a), such that the ribbon is strained, and may therefore become wrinkled or buckled at certain values of \u03b7 and T (Fig. 2b,c,g,h) [4, 5]. A second approach to the ribbon limit, initiated by Sadowsky [6] and revived recently by Korte et al. [7], considers the ribbon as an \u201cinextensible\u201d strip, whose shape is close to a creased helicoid\u2014an isometric (i.e., strainless) map of the unstretched, untwisted ribbon (Fig. 2d). A third approach, which may be valid for sufficiently small twist, assumes that the stretched-twisted ribbon is similar to the wrinkled shape of a planar, purely stretched rect- 1Our convention in this paper is to normalize lengths by the ribbon width W, and stresses by the stretching modulus Y, which is the product of the Young modulus and the ribbon thickness (non-italicized fonts are used for dimensional parameters and italicized fonts for dimensionless parameters). Thus, the actual thickness and length of the ribbon are, respectively, t = t \u00b7 W and L = L \u00b7 W, the actual force that pulls on the short edges is T \u00b7 YW, and the actual tension due to this pulling force is T = T \u00b7 Y. 139 Reprinted from the journal angular sheet, with a wrinkle\u2019s wavelength that vanishes as t \u2192 0 and increases with L [8]. Finally, considering the ribbon as a rod with highly anisotropic cross section, one may approach the problem by solving the Kirchoff\u2019s rod equations and carrying out stability analysis of the solution, obtaining unstable modes that resemble the looped shape (Fig. 2e) [9]. A recent experiment [3], which we briefly describe in Sect. 1.2, revealed some of the predicted patterns and indicated the validity of the corresponding theoretical approaches at certain regimes of the parameter plane (T , \u03b7) (Fig. 2). Motivated by this development, we introduce in this paper a unifying framework that clarifies the hidden assumptions underlying each theoretical approach, and identifies its validity range in the (T , \u03b7) plane for given values of t and L. Specifically, we show that a single theory, based on a covariant form of the F\u00f6ppl\u2013von K\u00e1rm\u00e1n (FvK) equations of elastic sheets, describes the parameter space (T , \u03b7, t,L\u22121) of a stretched twisted ribbon where all parameters in Eqs. (1) and (2) are assumed small",
            " Various \u201ccorners\u201d of this 4D parameter space are described by distinct singular limits of the governing equations of this theory, which yield qualitatively different types of patterns. This realization is illustrated in Fig. 3, which depicts the projection of the 4D parameter space on the (T ,\u03b7) plane, and indicates several regimes that are governed by different types of asymptotic expansions. 1.2 Experimental Observations The authors of [3] used Mylar ribbons, subjected them to various levels of tensile load and twist, and recorded the observed patterns in the parameter plane (T ,\u03b7), which we reproduce in Fig. 2. The experimental results indicate the existence of three major regimes that meet at a \u201c\u03bb-point\u201d (T\u03bb, \u03b7\u03bb). We describe below the morphology in each of the three regimes and the behavior of the curves that separate them: \u2022 The helicoidal shape (Fig. 2a) is observed if the twist \u03b7 is sufficiently small. For T < T\u03bb, the helicoid is observed for \u03b7 < \u03b7lon, where \u03b7lon \u2248 \u221a 24T is nearly independent on the ribbon thickness t . For T > T\u03bb, the helicoid is observed for \u03b7 < \u03b7tr, where \u03b7tr exhibits a strong dependence on the thickness (\u03b7tr \u223c \u221a t ) and a weak (or none) dependence on the Reprinted from the journal 140 tension T . The qualitative change at the \u03bb-point reflects two sharply different mechanisms by which the helicoidal shape becomes unstable. \u2022 As the twist exceeds \u03b7lon (for T < T\u03bb), the ribbon develops longitudinal wrinkles in a narrow zone around its centerline (Fig. 2b,c). Observations that are made close to the emergence of this wrinkle pattern revealed that both the wrinkle\u2019s wavelength and the width of the wrinkled zone scale as \u223c(t/ \u221a T )1/2. This observation is in excellent agreement with Green\u2019s characterization of the helicoidal state, based on the familiar FvK equations of elastic sheets [2]. Green\u2019s solution shows that the longitudinal stress at the helicoidal state becomes compressive around the ribbon centerline if \u03b7 > \u221a 24T , and the linear stability 141 Reprinted from the journal analysis of Coman and Bassom [5] yields the unstable wrinkling mode that relaxes the longitudinal compression. \u2022 As the twist exceeds \u03b7tr (for T > T\u03bb), the ribbon becomes buckled in the transverse direction (Fig. 2g), indicating the existence of transverse compression at the helicoidal state that increases with \u03b7. A transverse instability cannot be explained by Green\u2019s calculation, which yields no transverse stress [2], but has been predicted by Mockensturm [4], who studied the stability of the helicoidal state using the full nonlinear elasticity equations. Alas, Mockensturm\u2019s results were only numerical and did not reveal the scaling behavior \u03b7tr \u223c \u221a t observed in [3]. Furthermore, the nonlinear elasticity equations in [4] account for the inevitable geometric effect (large deflection of the twisted ribbon from its flat state), as well as a mechanical effect (non-Hookean stress-strain relation), whereas only the geometric effect seems to be relevant for the experimental conditions of [3]. \u2022 Turning back to T < T\u03bb, the ribbon exhibits two striking features as the twist \u03b7 is increased above the threshold value \u03b7lon. First, the longitudinally-wrinkled ribbon transforms to a shape that resembles the creased helicoid state predicted by [7] (Fig. 2d); this transformation becomes more prominent at small tension (i.e., decreasing T at a fixed value of \u03b7). Second, the ribbon undergoes a sharp, secondary transition, described in [3] as similar to the \u201clooping\u201d transition of rods [9\u201312] (Fig. 2e). At a given tension T < T\u03bb, this secondary instability occurs at a critical twist value that decreases with T , but is nevertheless significantly larger than \u03b7lon \u2248 \u221a 24T . \u2022 Finally, the parameter regime in the (T , \u03b7) plane bounded from below by this sec- ondary instability (for T < T\u03bb) and by the transverse buckling instability (for T > T\u03bb), is characterized by self-contact zones along the ribbon (Fig. 2e). The formation of loops (for T < T\u03bb) is found to be hysteretic unlike the transverse buckling instability (for T > T\u03bb). In a recent commentary [13], Santangelo recognized the challenge and the opportunity introduced to us by this experiment: \u201cAbove all, this paper is a challenge to theorists. Here, we have an experimental system that exhibits a wealth of morphological behavior as a function of a few parameters. Is there anything that can be said beyond the linear stability analysis of a uniform state",
            " (ii) A far-from-threshold (FT) expansion of the cFvK equations that describes the state of the ribbon when the twist exceeds the threshold value \u03b7lon for the longitudinal wrinkling instability. (iii) A new, asymptotic isometry equation (Eq. (42)), that describes the elastic energies of admissible states of the ribbon in the vicinity of the vertical axis in the parameter plane (T , \u03b7). We use the notion of \u201casymptotic isometry\u201d to indicate the unique nature by which the ribbon shape approaches the singular limit of vanishing thickness and tension (t \u2192 0, T \u2192 0 and fixed \u03b7 and L). We commence our study in Sect. 2 with the helicoidal state of the ribbon (Fig. 2a)\u2014 a highly symmetric state whose mechanics was addressed by Green through the standard FvK equations [2], which is valid for describing small deviations of an elastic sheet from its planar state. We employ a covariant form of the FvK theory for Hookean sheets (cFvK equations), which takes into full consideration the large deflection of the helicoidal shape from planarity. Our analysis of the cFvK equations provides an answer to question (A) above, curing a central shortcoming of Green\u2019s approach, which provides the longitudinal stress but predicts a vanishing transverse stress",
            " This result reflects the remarkable geometrical nature of the FT-longitudinally-wrinkled state, which becomes infinitely close to an isometric (i.e., strainless) map of a ribbon under finite twist \u03b7, in the singular limit t, T \u2192 0. At the singular hyper-plane (t = 0, T = 0), which corresponds to an ideal ribbon with no bending resistance and no exerted tension, the FT-longitudinally-wrinkled state is energetically equivalent to simpler, twist-accommodating isometries of the ribbon: the cylindrical shape (Fig. 5) and the creased helicoid shape (Fig. 2d, [7]). We argue that this degeneracy is removed in an infinitesimal neighborhood of the singular hyper-plane (i.e., t > 0, T > 0), where the energy of each asymptotically isometric state is described by a linear function of T with a t -independent slope and a t -dependent intercept. Specifically: Uj(t, T ) = AjT + Bj t 2\u03b2j , (42) 153 Reprinted from the journal where j labels the asymptotic isometry type (cylindrical, creased helicoid, longitudinal wrinkles), and 0 < \u03b2j < 1. For a fixed twist \u03b7 1, we argue that the intercept (Bt2\u03b2 ) is smallest for the cylindrical state, whereas the slope (A) is smallest for the FT-longitudinallywrinkled state",
            " In Table 2, we compare the control parameters and their relevant mutual ratios in both experiments. Green, who used a material with very large Young\u2019s modulus, could address the \u201cultra-low\u201d tension regime, T \u223c Tsm(t) (Fig. 3c), but a simple steel may exhibit a non-Hookean (or even inelastic) response at rather small T , which limits its usage for addressing the regime around and above the triple point (i.e., T > T\u03bb). In contrast, the experiment of [3] used a material with much lower Young\u2019s modulus, which allows investigation of the ribbon patterns in Fig. 2g, but the minimal exerted tension Tmin (associated with the experimental set-up) was not sufficiently small to probe Green\u2019s threshold plateau \u03b7lon(T ) \u2192 10t for T Tsm(t). This comparison reveals the basic difficulty in building a single set-up that exhibits clearly the whole plethora of shapes shown in Fig. 3. In addition to the effect of Tmin and THook, there is an obvious restriction on Lmax (at most few meters in a typical laboratory). Below we propose a couple of other materials, whose study\u2014through experiment and numerical simulations\u2014may enable a broader range of the ratios Tmin/Tsm, THook/T\u03bb and Lmaxt ",
            " The far-from-threshold analysis of the cFvK equations revealed a profound feature of the wrinkling instability: assuming a fixed twist \u03b7, and reducing the exerted tension (along a horizontal line in Fig. 3), the formation of longitudinal wrinkles that decorate the helicoidal shape enables a continuous, gradual relaxation of the elastic stress from the strained helicoidal shape at T > \u03b72/24, to an asymptotically strainless state at T \u2192 0. This remarkable feature led us to propose a general form of the asymptotic isometry equation (42), which characterizes the wrinkled state of the ribbon (Fig. 2b, c), as well as other admissible states at the limit T \u2192 0, such as the cylindrical wrapping (Fig. 2e) and the creased helicoid state (Fig. 2d). The asymptotic isometry equation provides a simple framework, in which the transitions between those morphologies in the vicinity of the vertical line (T = 0 in Fig. 3) correspond to the intersection points between linear functions of T (Fig. 6b), whose intercepts and slopes are determined solely by the geometry of each state. Beyond its role for the mechanics and morphological instabilities of ribbons, the asymptotic isometry equation may provide a valuable tool for studying the energetically favorable configurations of elastic sheets"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002355_tec.2020.3001914-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002355_tec.2020.3001914-Figure1-1.png",
        "caption": "Fig. 1. Cross-section of one pole of the PMSM under study.",
        "texts": [
            " The resulting surrogate functions as follows: it passes the inputs through the GPR to obtain PCA coefficients (or the scalar output), which are then passed through the inverse PCA transform to reconstruct the output waveform. We then use the PMSM surrogate to perform UQ and SA. Section III presents a case study. Specifically, we consider the effects of uncertainty on average torque, sixth torque harmonic magnitude, peak line-to-line voltage, and total core loss. The device under study is an eight-pole PMSM, rated for 235 Nm and 96 kW at 3900 rpm. Its cross-section is depicted in Fig. 1, where \u03b8rm is the mechanical rotor angle. The rotor lamination stack is divided into four identical sub-stacks, such that the rotor stack length equals the stator stack length, `m = 90 mm. During assembly, two of the four sub-stacks are rotated clockwise, whereas the other two are rotated counterclockwise, as shown in Fig. 2. A skew angle, \u03b8sk, is defined to parameterize this rotation. Note that Fig. 1 shows the cross-section of a sub-stack with a positive skew angle. It should be noted that the qd-axes are representative of an average magnetization direction from the combined effect of all four sub-stacks. To analyze sub-stacks, first we develop a parametric two-dimensional FEM. This high-fidelity FEM provides input-output observations that are used to construct a computationally inexpensive surrogate model that eventually replaces the FEM. To account for the skewing of the rotor, we combine two surrogate model results, obtained for positive and negative skew angles, respectively",
            "org/publications_standards/publications/rights/index.html for more information. the PM material is the uncertainty about the residual flux density, Br, which can be traced to the PM manufacturing process and the quality of the magnet material, among other factors [13], [14]. On the other hand, the representation of the uncertain magnetic properties of steel requires a more sophisticated model of the B-H characteristic and its spacial variability within the machine. The stator and rotor regions are divided into two subregions, as depicted in Fig. 1: (i) the degraded steel regions (dark shaded) close to the edges of the lamination, and (ii) the non-degraded steel regions (lightly shaded). The non-degraded steel properties may vary due to the grain structure uncertainty of the material, caused by the manufacturing process of nonoriented electrical steel sheets [15], [16]. The degraded steel regions model the effects of punching, where the degraded material proprieties depend primarily on two variables: (i) the distance from the cut edge and (ii) the magnetic field strength"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001717_j.addma.2020.101376-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001717_j.addma.2020.101376-Figure10-1.png",
        "caption": "Fig. 10. Sketch (a) and photos (b, c) of the experimental setup. To make the resin level consistent for each measurement the initial recoating at a low speed (5 mm/s) was implemented (move \u201c1\u201d in (a) and (b)). After initial recoating the level of the resin is measured along the centerline of the cavity (dashed line in (a)). Immediately after that the actual recoating step takes place (move \u201c2\u201d in (a) and (c)) followed by the surface height profile measurement.",
        "texts": [
            " It has been found that the non-uniformity of the resin surface under the deep cavity levels due to gravity relatively quickly (< 2 s). The fluid physics of the development of the leading edge bulge during the deep-toshallow transition of the underlying geometry has been described. The influence of the recoater speed on the resin level above the cavity also showed a good agreement with the theoretically predicted values. A verification of the numerical calculations is needed in order to establish the usefulness of 2D simulations and its ability to generate complementary data. The experimental setup shown in Fig. 10 is used to reproduce the recoating step in a real printing process. A rectangular metal block with a rectangular-shaped cavity in it is used to mimic the non-uniform underlying geometry. A metal bar with a rectangular cross-section mounted on the moving stage serves as the recoating blade. In the experiments, measurements of the resin surface topography have been accomplished by a point-wise distance sensor based on the confocal chromatic technique with the specifications presented in Table 2. The confocal chromatic technique is based on the phenomenon of chromatic dispersion in which the refractive index depends on the wavelength of the light"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002765_s11071-021-06327-0-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002765_s11071-021-06327-0-Figure14-1.png",
        "caption": "Fig. 14 Extra-pitch meshing",
        "texts": [
            " There is a time difference between tangential and radial impact forces for the sprocket. This time delay allows the impact energy to be distributed over a longer time interval, Fig. 10 Meshing tooth number variation in equalpitch meshing thereby reducing track vibration. For this reason, it gets the least bottom contact force in three cases. In contrast to sub-pitch meshing, traction is transmitted by the tooth that enters the mesh at the latest. Along the clockwise direction, regular intervals D, 2D, 3D, 4D appear between the engaged pins and left surfaces, as highlighted in Fig. 14. Similarly, in this case, the performance of the track and sprocket depends on the value of D. When D\\PC/4, the crawler can still mesh normally, but scraping may occur between pin edges and teeth. Figure 15 details the meshing tooth number variation. Different from equal-pitch meshing, there is no regular periodic change. This number is mainly 4 or 5, but 2 and 3 also occur. When D\\PC/4, the number of meshing teeth will decrease at each moment. Further, \u2018\u2018stuck in teeth\u2019\u2019 may occur, resulting in the failure of crawler to operate properly similarly"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure40.3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure40.3-1.png",
        "caption": "Fig. 40.3 Stress test results",
        "texts": [
            " However, graphite plastic is not suitable, since its dielectric properties remain in question. On high-voltage lines, it will become a conductor and create a short circuit between the robot and thewire.Nylon is not used, as it has a lowcoefficient of friction, which will not allow for movement along a wire that has sagging. Comparisons of various materials were carried out taking into account the loads on the wheel under the action of gravity, torque and crosswind. All the materials examined showed a good result in wheel deformation (Fig. 40.3) and safety margin. The main stresses are on the axis of rotation of the wheel and the contact surface with the wire. In the future, when using other materials, you can remove excess material from the outer edges of the wheels. Further studies were conducted to determine the basic material for the robot body. Initially, the form was to be in the form of the Latin letter \u201cG\u201d. This shape allows you to manually hang the robot on a cable, hanging it from the side of the wire, and had a streamlined shape that allows you to protect internal mechanisms and electronics from weather conditions with minimal junction points"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001952_j.mechmachtheory.2020.104125-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001952_j.mechmachtheory.2020.104125-Figure1-1.png",
        "caption": "Fig. 1. A 7-link 2-DOF PGT and its conventional graph.",
        "texts": [
            " The synthesis results of 8- and 9-link 2-DOF PGTs, to the best of our knowledge, are new results that have not been re- ported in literature. The derived atlas of non-fractionated 2-DOF PGTs enables designers to select desired PGTs for target application. The first graph representation called conventional graph of PGT was presented by Buchsbaum and Freudenstein [8] . A link of PGT is denoted by a vertex and a kinematic pair is denoted by an edge. Revolute and geared pairs are distinguished by different types of edges. For example, Fig. 1 (a) shows the functional diagram of a 7-link 2-DOF PGT and Fig. 1 (b) shows its conventional graph. In this paper, revolute and geared pairs are denoted by solid and dashed edges, respectively. Letters a, b, c and d are used to distinguish different levels (positions) of axes of rotation. All the kinematic pairs in a conventional graph are assumed as simple (binary) joint. However, it is well known that the same-level revolute pairs are equivalent to a multiple joint [14 , 15] . Hus and his coworkers [14 , 15] developed a kind of graph representation of PGTs by applying a solid polygon to denote a multiple joint",
            " [20 , 21] presented a new kind of graph representation where a multiple joint is denoted by a hollow vertex. According to the definition in Refs. [20 , 21] , a rotation graph of PGT is acquired as follows. Vertices of each geared edge are directly connected to the associated transfer vertex by revolute edges. If there exists a loop (or loops) formed exclusively by revolute edges, vertices in such a loop (or loops) are connected to a hollow vertex representing a multiple joint. For example, the rotation graph of Fig. 1 (b) is illustrated in Fig. 2 (a) where vertex 8 denotes a multiple joint. When the vertices of the same-level edges in a rotation graph are connected to a hollow vertex, the corresponding displacement graph is generated [20 , 21] . For example, the displacement graph of Fig. 1 (b) is shown in Fig. 2 (b). A d -degree hollow vertex is equivalent to d - 1 revolute pairs. Both the adjacency matrices of the rotation graph and displacement graph are defined as follows: [ a i, j ] n \u00d7n = \u23a7 \u23aa \u23a8 \u23aa \u23a9 1 , if vertex i is connected to vertex j by a solid edge 2 , if vertex i is connected to vertex j by a dashed edge d, if i = j and i is the label of a hollow vertex 0 , otherwise where n is the number of vertices on the graph and d is the degree of a hollow vertex. For example, the adjacency matrix of the displacement graph in Fig. 2 (b) is shown in Fig. 2 (c). If all the edges in a conventional graph of PGT are assumed to be revolute edges, the derived graph is its parent graph. An element a i , j of the adjacency matrix equals 1 if vertices i and j are adjacent; otherwise, it equals 0. For example, the parent graph of Fig. 1 (b) is shown in Fig. 3 (a) and its adjacency matrix is shown in Fig. 3 (b). If a graph can be separated into two independent sub-graphs at a vertex, this vertex is called a cut vertex and this graph is said to be vertex-fractionated. For example, Fig. 4 (a) shows a vertex-fractionated parent graph where vertex 3 is a cut vertex. If a graph can be separated into two independent sub-graphs at an edge, this edge is called a bridge and this graph is said to be edge-fractionated. For exam- ple, Fig. 4 (b) shows an edge-fractionated parent graph where edge e 15 is a bridge"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000281_iecon.2019.8927827-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000281_iecon.2019.8927827-Figure14-1.png",
        "caption": "Fig. 14 \u2013 Assembly of the considered PM synchronous generator: rotor (a) and evaluation of magnetic induction in airgap (b)",
        "texts": [
            " The bonded magnets help the implementation of that high number of slices with respect to the sintered magnets, thus allowing the realization of the Halbach pattern. It is important to note that Halbach and Epoxy Bonded magnets have the same magnetic characteristics; the difference is represented by the magnetisation pattern. NdFeB bonded magnets characteristic places them between sintered NdFeB materials and ferrites; those properties make them suited for the substitution of NdFeB sintered magnets (Fig. 13), especially in over-fluxed machines. The bonded magnets were applied with good results to a three-phase PM synchronous generator prototype (Fig. 14); it resulted in a slight increase of the stator cost due to the need of several additional turns in the stator windings and a consistent reduction of the rotor cost due to the change of magnetic material technology. The output voltages of the machines equipped with the two considered magnets are identical (Fig. 15). Where compact geometries and weight reduction become a challenging constraint, the powder material technology can be considered, both for the stator soft magnetic structures (in general, Soft Magnetic Composite are adopted), and for the permanent magnets, allowing different freedom degrees to the machine designers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002541_0954407020974497-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002541_0954407020974497-Figure2-1.png",
        "caption": "Figure 2. Ackerman steering wheel mechanism: (a) condition for Ackerman steering and (b) % Ackerman measurement criteria.",
        "texts": [
            " For smooth turning of a vehicle, all four wheels should follow arcs having a common center, known as instantaneous center (IC). The Ackerman steering geometry intends to have such performance during cornering so as to avoid the tyres scrubbing or slip sideward. Therefore, the configuration of four wheels during turning is such that the IC of the front wheels meets the IC of the rear wheels is said to be a True Ackerman criteria. In race cars, which are front wheel steered vehicles, this IC will lie on the projection of the rear axle during a low-speed turn (Figure 2(a)). For Ackerman steering mechanism of a vehicle having track width (w) and wheel base (l), the relation between inner and outer wheel steering angles can be written as1: ui =tan 1 l l tan uo w ! \u00f01\u00de Equation (2) shows percentage Ackerman, which is a ratio between the actual steer angle to that angle correspond to zero scrub. Percentage Ackerman is a criterion for measuring the deviation of actual steer angle from true Ackerman geometry. %Ackerman= ui uo uA i uA o 3100 \u00f02\u00de Where uAi and uA o are inner and outer wheel steering angles respectively with respect to true Ackerman criteria. According to property of IC, each wheel\u2019s direction of motion is normal to the line drawn from the wheel to the instantaneous centre I, as shown in Figure 2(b). Basically the inner wheel steer angle may be greater or equal to outer wheel steer angle, that is, ui=uo\u00f0 \u00de\u00f8 1. The steer angle ratio can never be one in a kinematic chain other than a linkage which forms a parallelogram. In other words, the curvature of inner wheel is larger than outer wheel, since they negotiate turns of different radii. Therefore, the increasing ratio is always anticipated. The relation of actual and desired angle of rotation of the wheel are written as, ui = uA i us \u00f03a\u00de uo = uA o + us \u00f03b\u00de Here us can be termed as the error or the angular deviation from true Ackerman",
            " The lower limit for the wheelbase is fixed to 60 inches or 1525 mm, as mentioned in the Rule book.2 Since the tie arm is built into the knuckle, its length (a) should be such, so as to accommodate within the interior of the wheel or rim. A minimum length of tie arm is decided by the minimum steering effort that would turn a wheel about the king pin axis or steering axis. Moreover, the upper bound for the parameter a is decided by the space available inside the rim in addition to the steering effort. As Figure 2 shows that the front axle steering kinematics along with the wheel assembly; it is clear that if the angle of inclination of tie arm with respect to the rack (finitial) is beyond a particular value, there will be contact between the tie rod and the inner surface of the rim during full tyre deflection. Considering the wheel kinematics, it was observed that for the value of finitial beyond 170 degrees will result in hindrance between the rim and the tie rod. This estimation also puts the limit for the length of the tie rod (b) in the range (200\u2013450 mm)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001208_j.triboint.2019.105999-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001208_j.triboint.2019.105999-Figure6-1.png",
        "caption": "Fig. 6. Internal cavity hydraulic oil compensation seal.",
        "texts": [
            " The oil pressure is used to improve the sealing performance of the piston rod seal by introducing external oil or using internal cavity oil to act on a special structure of the seal, which is called a pressure compensation seal. For example, Zhou and Jiang proposed a pressure compensation sealing technology, which introduces the high pressure oil of the hydraulic cylinder into the action groove of the special sealing ring to increase the contact pressure of the sealing pair and compensate for the seal wear [7], as shown in Fig. 6. Hunger Company developed an intelligent EVD (Extern Vorspannbare Dicht system) sealing device. The device has an internal Nomenclature D rod diameter, m PRMS1 contact pressure RMS of external stroke, MPa E elastic modulus, GPa FN, F\u2019N contact force of seal pair, MPa PAmin minimum contact pressure of external stroke static seal, MPa Ff1, Ff2 axial friction of seal pair, MPa fRMS friction force RMS in the external stroke, MPa PARMS contact pressure RMS of external stroke static seal, MPa H displacement of piston rod external stroke, m PA contact pressure of LA, MPa Sx(f)max frictional power spectrum peak ho oil film thickness, mm V surface speed of rod, m/s havg average film thickness, mm V0 leakage of external stroke, ml hRMS oil film thickness RMS, mm WA maximum pressure gradient L contact length of sealing area, mm WE wavelet packet energy entropy LA the point with the highest pressure gradient \u03c5 Poisson\u2019s ratio \u03b2 the lowest energy ratio P fluid pressure, MPa \u03c1 density of the liquid,kg/m3 P(x) contact pressure of seal pair, MPa \u03b7 dynamic viscosity,Pa* s P0 compensating pressure, MPa \u03c4\u00f0x\u00de friction stress of sealing pair, MPa P1 pressure of the test cylinder, MPa \u03c4  max maximum friction stress, MPa Pmax maximum contact pressure of seal pair, MPa Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001502_tia.2020.2972835-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001502_tia.2020.2972835-Figure2-1.png",
        "caption": "Fig. 2. Cross section of a SAG mill.",
        "texts": [
            " 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/TIE.2014.2378732 synchronous ring motor. The stator is separately mounted to the basement. It is fed by a load-commutated cycloconverter with a power rating of typically 20 MW at a maximum fundamental frequency of 6 Hz [3], [4]. The converter is equipped with thyristors, which are the most powerful semiconductor devices available. The interior of a SAG mill is shown as a cross section in Fig. 2 [5]. The shaded area in the lower right portion indicates 0278-0046 \u00a9 2014 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. where the grinding material and a number of steel balls are encountered. The cylinder rotates in the anticlockwise direction, which elevates the steel balls and the major-sized lumps of ungrounded material up to a certain level"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000471_j.triboint.2016.02.019-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000471_j.triboint.2016.02.019-Figure3-1.png",
        "caption": "Fig. 3. Schematics of (a) macroscale tribometer (CETR-UMT-2, Bruker), (b) and sample configuration. To fix wiper and label samples, they were mounted onto polypropylene (PP) stationary upper cylinder and rotating lower PP disk, respectively.",
        "texts": [
            " Scan distance of 500 mm, which is shorter than the width of wear track, were taken in order to characterize the surface roughness of the worn area. Sampling rate was 1 datum point per 0.25 mm. All measurements were performed at four different locations on each sample in order to obtain root mean square (RMS) roughness which is commonly used as a measure of roughness [17]. Table 2 shows RMS roughness data of label and wiper samples. A commercial friction tester (CETR-UMT-2, Bruker) was used to measure static and kinetic friction (Fig. 3). A standard cylinder-onflat geometry configuration was used [17]. The wiper and label samples were mounted onto a polypropylene (PP) cylinder that was 19.1 mm in diameter and 6.35 mm thick, and a PP disk that was 76.2 mm in diameter and 3.18 mm thick, respectively. The reason for selection of PP as a substrate of wiper and label samples is that mechanical properties of PP are similar to a real wiper substrate generally made of PU, and a container generally made of PP. In the configuration selected, the wiper remains stationary and contacts the label sample during sliding test at a normal load corresponding to desired contact pressure as in a real labeling system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002009_s42417-020-00259-6-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002009_s42417-020-00259-6-Figure2-1.png",
        "caption": "Fig. 2 Planetary gear set and system dynamics model",
        "texts": [
            " Finally, the stress time history of the ring gear under the dynamic load is solved step by step using the Newmark-\u03b2 time integration method. To verify the effectiveness of the above calculation method, the time-domain strain history of the root of the ring gear in the planetary gear set of the wind turbine test bed is extracted using an FBG sensor and is then compared with the theoretical calculation results. Basic parameters of planetary gear transmission system are shown in Table\u00a01. The model of planetary gear transmission system is shown in Fig.\u00a02. The planetary gear system is 2K-H planetary gear set, and the four planets are evenly distributed along the circumference. In the process of planetary gear transmission, the motion process is more complicated than that of fixed-shaft transmission. Also, the analysis is much more difficult. Therefore, considering the complex coupling relationship among various components, the translation\u2013torsion coupling dynamic model of the planetary gear transmission system is established. In the modeling, the center of the model is taken as the coordinate origin, the horizontal direction is X, and the vertical direction is Y"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003910_s1474-6670(17)45681-3-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003910_s1474-6670(17)45681-3-Figure2-1.png",
        "caption": "Figure 2: Clarification of the piston shaft displacement in a torque sensor",
        "texts": [
            " The two moving sheaves are connected to two corresponding hydraulic cylinders and pistons. The primary pulley is regarded as the drive end (input end) and the second ary pulley as the driven end (output end) of the CVT. The components special in this CVT are the so-called primary and secondary torque sensors, and the control unit. A torque at the input of a torque sensor causes an axial sliding of its piston shaft which is enabled by a rolling of the balls in the torque sensor (Figure I). This displacement of the piston shaft changes the drain orifice as depicted in Figure 2. Depending on the drain orifice area and the fluid flow into the torque sensor, a pressure results which is a measure for the torque transferred by the CVT. As shown in Figure I. the two torque sensors are hydrau- IicallY connected in series. For two identical torque sensors, the sensor which measures the higher torque determines the pressure in the cylinder. This double torque sensor system prevents a slipping of the belt at the critical secondary pulley by produc ing the nearly optimal steady state pressure [7]",
            " In this case the torque sensor represents a rigid connection. By increasing the torque or reducing the flow, a rela tive motion between the input and the output shaft of the torque sensor is possible. Now the torque sensor represents an elastic connection, and the torques M I and M2 are determined by the pressure in the torque sensors as well as by the friction torque MR as fol lows: (I) where cF is a characteristic quantity of the torque sensor which relates the rotation of the shaft to the translatory motion of the piston ring (Figure 2). A change from an elastic to a rigid connection or vice versa causes a change in the structure of the mechani cal scheme. Since the simulation software cannot deal with structural changes, a stiff spring-damper element is used instead of a rigid connection. In this case the torques M ( and M2 are given as follows: M( = kMF6q>( + CMF60>1 M2 = kMF6q>1I + CMF60>1I . (3) (4) With these two torques we obtain the following equa tions of motion for the input and the output shaft of theCVT: eac.Oa = Ma - M( ezc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003646_s0109-5641(99)00031-7-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003646_s0109-5641(99)00031-7-Figure1-1.png",
        "caption": "Fig. 1. The dimensional accuracy of a cast crown is expressed by the discrepancy (x 2 y) measured on its wax pattern and the casting on the same metal die.",
        "texts": [
            " Overshooting of up to about 58C from the target temperature occurred during the changeover of the heating programme to the cooling one. Five specimens were prepared from two different mixes for each material. TE values determined at various temperatures were expressed as a percentage against the original height of specimen. The results from investment A were subjected to two sample Student\u2019s t-test assuming unequal variances for different batches and TE measuring stresses. The metal die method previously described [2] was followed for the assessment of full crown accuracy. This is schematically shown in Fig. 1. Full crown wax patterns still seated on metal dies were annealed at room temperature with sprues attached for 1 h. The pattern and then a spacer were removed and the former returned to the die using an index line engraved on the pattern which aligned with a line on the die. The distance separating the gingival margin of the wax crown and the shoulder of the metal die (x), was measured at four fixed points, 908 apart, around the die. The measured values were close to the thickness of the spacer",
            " All castings were carefully removed from the mould, scrubbed under running water, and cleaned in water using an ultrasonic cleaner. Each casting was finally cleaned individually in a solution (1% HF 1 13% HNO3) for 10 min again using the ultrasonic cleaner. The castings were carefully examined for any irregularity and nodules and a cutting disc was used to separate the crown from the sprue. Each crown was returned to the original metal die, seated under a load of 20 N and the distance separating the gingival margin of the crown and the shoulder of the metal die (y) (Fig. 1) was measured at the same four fixed points as for the wax pattern. The dimensional accuracy was expressed by the discrepancy of the two measurements (x 2 y) for each point and the mean was calculated from the four measurements for each crown. Ambient laboratory conditions were 22 ^ 18C and 50 ^ 10% relative humidity. The simple least squares linear regression analysis was made using all accuracy data collected from each crown against a mean TE value of each investment. The equation is Y b0 1 b1X, where Y is crown accuracy expressed by the discrepancy (x 2 y) in mm and X is investment TE in %"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001041_s40998-019-00214-6-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001041_s40998-019-00214-6-Figure1-1.png",
        "caption": "Fig. 1 Frames representation",
        "texts": [
            " Section\u00a04 details and discusses the simulation and experimental results using robotic operating system (ROS) implemented on a commercial platform named AR.Drone V2. Section\u00a05 gives some conclusions and future work. To test the proposed intelligent autopilot, a simulation environment is developed based on the nonlinear mathematical model of the Quadrotor. The mathematic model [already developed in Zareb et\u00a0al. (2013) based on the work of Beard (2008)] contains the equations of motion, using the derivation of Coriolis, in which we use the following coordinate frames (Fig.\u00a01): the inertial frame, the vehicle frame, and the body frame. The notation and coordinate frames are typical 1 3 in the aeronautics literature. The kinematic and dynamic equations are presented. The Quadrotor\u2019s model can be written as: with s = sin( ) and c = cos( ). The same notation applies for s , c , s and c . where x, y and z are the inertial position of the Quadrotor along xi , yi and \u2212zi in Ri (inertial frame). (u, v and w) are the body frame velocity measured along xv , yv and zv in Rv (vehicle frame)",
            " The GA \u2212 Outer evolutionary process is presented in the following pseudo-code: (42)f2 = 1\u2215(tf \u2212 ti)\u222b tf ti (e2 4 (t) + e2 5 (t))d , 1 3 To find the optimal values of the PID-fuzzy controllers P\u2217 2 = (K e ,K de ,K pd ,K pi ) the process followed in the GAOutput is the same as in the GA-Inner, the difference is in the number of the variables, which is eight parameters and the modes of flight. At the end of the EIFC approach, the optimal parameters of IFC are P\u2217 = [P\u2217 1 ,P\u2217 2 ] . In order to prove the effectiveness of the EIFC approach, simulation and experiment tests are conducted in the next section. To validate the EIFC approach, first simulation tests are made on AR.Drone V2 Quadrotor shown in Fig.\u00a01. In fact, it was defined as researcher platform in Tom\u00e1\u0161 et\u00a0al. (2011) which is one of the reasons to choose it, and the second reason is its low cost. Based on the nonlinear model and the physical parameters (in Table\u00a01) and the GA parameters shown in Table\u00a02, the off-line EIFC approach calculates the optimal parameters P\u2217 1 and P\u2217 2 on MATLAB-Simulink of the IFC of the AR.Drone V2. \u00a0Table\u00a03 summarizes these values. The choice of the initial generation has an impact on the convergence of GA. For these reasons, the choice is made to take the value for proportional and derivative bigger than integral factors for GA-Inner P1(0) ",
            " = l Jyy C\ud835\udf03 (52)?\u0308? = l Jzz C\ud835\udf13 where the EIFC parameters ensure the stability of the roll close-loop. Here, from the stability analyzes of the fastest dynamics (pith and the roll), the stability of the IFC control system is guaranteed. In the next subsection, the methodology to implement the EIFC parameters using the Robot operating System (ROS) on the AR.Drone is presented in detail. For the experiment, AR.Drone V2 is used; it is a commercial Micro-UAV Quadrotor name created by the French company Parrot (Fig.\u00a01). It is usable as research and education platform as in Tom\u00e1\u0161 et\u00a0al. (2011). It is equipped with several sensors to ensure basic tasks of position stabilization: 2-axis 1 3 Fig. 19 Roll angles comparison by using manual and Bi-GA configurations 40 45 50 55 0 2 4 6 8 time(sec) \u03c6 (d eg re e) Reference With Bi\u2212GA configuration With Manuel configuration 34 36 38 40 42 44 46 \u22120.5 0 0.5 1 1.5 2 2.5 3 time(sec) \u03c6 er ro r Manual configuration Bi\u2212GA configuration Fig. 20 Yaw angle comparison by using manual and Bi-GA configurations 5 10 15 20 25 30 5 6 7 8 9 10 11 time(sec) \u03c8 ( de gr ee ) Using Bi\u2212GA configuration Reference Using Manual configuration Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001152_ab3f56-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001152_ab3f56-Figure4-1.png",
        "caption": "Figure 4. Wire actuation system generating compression force.",
        "texts": [
            " The dielectric layer has protrusions in micro-scale to increase the sensitivity [16]. The dielectric layer can be easily deformed due to the air gap between the protrusions (figure 3(A)). The dielectric layer was fabricated by molding elastomer, and the fabric electrodes were bonded to the dielectric layer utilizing a thin uncured silicone layer through a spin-coating process (figure 3(B)). A micro metal gear motor (HPCB, Pololu, NV, USA) with a small size pulley was used to pull the wire that connects the ends of the sleeve\u2019s outer layers, as shown in figure 4. By controlling the degree of wire pulling, the motor can change the compression force of the sleeve. The housing to mount the gear motor was installed on a base plate integrated within the sleeve between the outer and inner layers. The base plate was made of a thermoplastic polyurethane (TPU) material. As this material is flexible, the base plate is deformed to fit the body when worn, providing comfort. The soft pressure sensor was installed underneath the base plate for detecting the compression force applied by the outer layer"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001447_s12555-018-0369-2-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001447_s12555-018-0369-2-Figure6-1.png",
        "caption": "Fig. 6. Diagrammatic sketch of varing LOS.",
        "texts": [
            " Earth angle b. Fig. 5. Push-broom velocity. D \ud835\udc97\ud835\udc91 L \ud835\udf36 \ud835\udc79\ud835\udc6c \ud835\udc79\ud835\udc6c S H O \ud835\udc76\ud835\udc72 \ud835\udf4e\ud835\udc91\ud835\udc73 Fig. 5. Push-broom velocity. where L is the slant range from the instantaneous imaging ground point D to the satellite, which is determined by L = (RE +H)cos\u03b1 \u2212 \u221a RE 2 \u2212 (RE +H)2sin2\u03b1. (23) Hence, the image motion velocity at the point D (the speed of LOS relative to imaging point on the ground) considering the Earth rotation as vi =vD + vp \u2212 vE = \u221a\u221a\u221a\u221a(vp + vD cos\u03b7 \u2212 vE cos(\u03b7 \u2212 i))2 +(vD sin\u03b7 \u2212 vE sin(\u03b7 \u2212 i))2. (24) Seen from Fig. 6, when the satellite is in push-broom mode during attitude maneuver, the LOS is varying all the time. Let D be the imaging target at a given time and D1 be the expected target in the next moment. However, due to the transport motion caused by the orbital motion and the Fig. 6. Diagrammatic sketch of varing LOS. Fig. 7. Drift angle for NPGTI mode. Fig. 7. Drift angle for NPGTI mode. migration caused by the Earth\u2019s rotation, the actual imag- ing point on the Earth surface in the next moment might be D2, namely the imaging point changes from D1 to D2, and this phenomenon should be compensated for through suitable attitude steering of the satellite. Thus, the drift angle is defined as the angle between the satellite projection scan velocity vp and the image motion velocity vi showed in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001810_s40436-020-00317-y-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001810_s40436-020-00317-y-Figure1-1.png",
        "caption": "Fig. 1 Structure diagram of equipment a and b the four-laser PBF equipment, c optical system in the developed PBF equipment",
        "texts": [
            " The online control of the temperature field during sintering was realized to improve the quality of the fabricated components. An online system for monitoring the sintering temperature field was integrated into the PBF equipment using an infrared thermal imager. Figures 1a and b show a four-laser PBF equipment that was designed and constructed with a working area of 2 000 mm 9 2 000 mm. It primarily comprised two systems: a lifting system and an optical system, and an infrared thermal imager was integrated into the optical system. As shown in Fig. 1c, the infrared thermal imager (Flir A310) was installed at the center of the optical system in the equipment. This installation method prevents the shape distortions of the shooting target and ensures the accuracy of the temperature measurement. A 90 9 73 wide-angle lens was used in the infrared thermal imager to capture the entire 2 000 mm 9 2 000 mm powder bed area, whereas the infrared thermal imager was nested inside a water-cooled protective cover to protect it from the high ambient temperature"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000204_j.autcon.2019.102996-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000204_j.autcon.2019.102996-Figure16-1.png",
        "caption": "Fig. 16. Internal and external cylinder are used as aids to create geometry \u2013 the internal cylinder is related to the hub's diameter, and the external cylinder is larger than the hub's diameter; subpart A is inside the internal cylinder; subpart B is between internal and external cylinder; subpart C is outside the external cylinder; and the extension vector determines the length of subpart C, it is aligned with elements' axes.",
        "texts": [
            " From this measurement, only, it is possible to calculate all other angles required, since notch vectors are at a constant angular distance. The correction angle for this example is 7.57\u00b0 counterclockwise (Fig. 15b). The same procedure applies, automatically, to any other hub with a different number of notches. After placing the hub in each mesh node, we can generate the adjusted geometry of the interconnecting parts. The interconnecting part geometry subdivides into three subparts: subpart A of the geometry depends on notch location; subpart C depends on element location, and subpart B connects A to C. Fig. 16 shows these subparts. 3.3.8.1. Generation of subpart A. Subpart A depends on the geometry and positioning of the hub notch. Fig. 17a shows the geometry of the hub notch with its outline and its direction vector. The outline of the notch may be modified according to the profile available. Next, the curve is extruded to form the surface of subpart A (Fig. 17b). The hub axis vector guides the extrusion of the subparts related to the corresponding hub. This step completes after top and bottom surfaces cap the notch geometry (Fig",
            " We simplified the circular and rectangular sections into segments of curves to avoid geometry errors in the program. Fig. 18a shows this simplification and subsequent surface generation. All quarters join into one surface; then, the joined surface connects to subparts A and C (Fig. 18b). 3.3.8.3. Generation of subpart C. The segments of mesh lines depicted in Fig. 12a are base for generation of subparts C, which are pipes with predefined radiuses. The start point of the pipe is the intersection of the segments of mesh lines and the external cylinder (Fig. 16). External and internal cylinders define a transition zone from subpart A to subpart C. The part between both cylinders is called subpart B. The radius of the internal cylinder is the same as the radius of the hub, and the radius of the external cylinder is larger than the internal one by any amount necessary for the transition. The pipe created at the defined intersection point extends through a vector of any amplitude desired, in the direction of the element centerline; it is the extension vector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000529_j.apm.2016.03.048-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000529_j.apm.2016.03.048-Figure6-1.png",
        "caption": "Fig. 6. Spherical squeeze film bearing (for \u03d5 o = \u03c0 2 ).",
        "texts": [
            "org/10.1016/j.apm.2016.03.048 8 A. Walicka et al. / Applied Mathematical Modelling 0 0 0 (2016) 1\u201312 Fig. 5 presents the load-carrying capacity \u02dc N as a function of the squeezing ratio \u025b . This load capacity is similarly induced by the surface roughness and the dimensionless coefficient of pseudo-plasticity \u03bb; the distributions of these mechanical parameters are similar to those presented in [19,20] . 6. Spherical squeeze film bearing Let us consider now a spherical squeeze film bearing shown in Fig. 6 . Introducing the following parameters: u = \u02dc h = h C = 1 \u2212 \u03b5 cos \u03d5, u o = 1 \u2212 \u03b5 cos \u03d5 o , \u2202 \u0303 h \u2202t = \u2212 \u02d9 \u03b5 cos \u03d5, e = 1 \u2212 \u03b5, \u03bb = k ( \u03bc \u02d9 \u03b5R r C )2 , \u02dc p = ( E p \u2212 p o ) C 2 \u03bc \u02d9 \u03b5R 2 r , \u02dc N = N C 2 \u03bc \u02d9 \u03b5R 4 r , (6.1) we will obtain the following formulae for pressure distribution and load-carrying capacity: \u02dc p = 6 \u03b5 [ F ( u o ) \u2212 F ( u ) ] , (6.2) \u02dc N = 6 \u03c0 \u03b5 3 {( e 2 \u2212 2 e ) [ F ( u o ) \u2212 F ( e ) ] + W ( u o ) \u2212 W ( e ) } , (6.3) where F ( u ) = I ( u ) \u2212 27 5 \u03b5 2 \u03bbJ ( u ) . (6.4) Functions I ( u ), J ( u ), W ( u ) are given in Appendix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003688_1.1304914-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003688_1.1304914-Figure4-1.png",
        "caption": "Fig. 4 Rolling contact between the rotor and track as viewed from the negative e2 direction",
        "texts": [
            " Given the relative rotation of the rotor, we use the methods of Casey and Lam @6# to calculate the angular velocity vector of the rotor relative to the track v\u0302r ,t : v\u0302r ,t5vr2vt5g\u0307e11a\u0307e3 (2) where vr is the absolute angular velocity vector of the rotor. In postulating constraints on the motion of the rotor, several conditions could exist at the contact point between the track and the rotor\u2019s axle. These include frictionless sliding, sliding with friction, and rolling without sliding. We shall only consider the last case because it is the only condition that presents a mechanism for spin-up of the rotor. Figure 4 shows the proposed type of contact between the rotor and the track. The axle contacts the track at a point P on the track\u2019s lower surface and at a point Q on the upper surface. In this configuration, the center of mass of the rotor remains coincident with the center of the track as the rotor rolls at both points P and Q. 000 by ASME JUNE 2000, Vol. 67 \u00d5 321 15 Terms of Use: http://asme.org/terms Downloaded F Introducing a right-handed orthonormal basis $ep1 ,ep2 ,ep3%, as illustrated in Fig. 4, facilitates the analysis. Here, ep1 points from the center of the rotor ~point O! toward point P such that the angle b between ep1 and e1 is b5tan21S Ra Rt D . (3) Furthermore, ep2 is parallel to e2 . Rolling at P and Q introduces the following constraint equations: ~vr1vr3pP!2~vt1vt3pP!50, (4) ~vr1vr3pQ!2~vt1vt3pQ!50, where vr is the velocity vector of the rotor\u2019s center of mass, vt is the velocity vector of the center of mass of the track, and pP and pQ are the position vectors of P and Q relative to the rotor\u2019s center of mass, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001981_1748006x20964614-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001981_1748006x20964614-Figure6-1.png",
        "caption": "Figure 6. Schematic and overview of the gearbox comprehensive test rig.",
        "texts": [
            " The training process of CANN is completed after multiple backpropagation training and fine tuning. The test samples are transformed into Hilbert envelope spectrum and normalized to the range of (0,1) with the equation (15), and then input them into the trained CANN to achieve gearbox fault diagnosis. In gearbox fault diagnosis, it is very important to construct sensitive features to improve the accuracy of fault identification. In this section, the gearbox fault simulation experiment is performed on the drivetrain diagnostic simulator (DDS). The test bench is illustrated in Figure 6. The test bench of DDS mainly consists of drive motor, transmission shaft, gearbox, loading system, DC speed regulating system, acceleration sensor, charge amplifier, DP/INV306U data acquisition instrument and industrial computer. The gearbox is composed of the two-stage planetary gearbox and the two-stage fixed shaft gearbox. The speed of the drive motor is set to 1200 r/min and the load of the magnetic powder brake is set to 5kg. During the experiment, four kinds of sun gear faults and three kinds of bearing faults are simulated with electrical discharge machining, including tooth missing fault, tooth surface wear, tooth chipped fault, tooth root crack, bearing outer race fault, inner race fault and rolling element fault, as described in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure6-1.png",
        "caption": "Fig. 6 3D model of circular arc slicing cutter with curved rake face",
        "texts": [
            " That is hi u;w\u00f0 \u00de \u00bc \u2211 H\u22121 i\u00bc0 \u2211 I\u22121 j\u00bc0 Qi; jQi;3 u\u00f0 \u00deQj;3 w\u00f0 \u00de 0\u2264u;w\u22641 \u00f015\u00de This section takes the circular arc tooth of the pinwheel housing in an RV40E reducer as an example to illustrate the feasibility of the cutter design method proposed in this paper. The parameters of the workpiece are shown in Table 1. The cutter parameters and processing parameters are set as shown in Table 2 and Table 3. Based on the above parameters, the cutting edge, the curved rake face, and the flank face are calculated by the method described in Sections 2.2 and 2.3. And the 3D model is generated as shown in Fig. 6. Different from design angles (or called shape angles), the working angles are dynamic in regard to the cutting velocity, which directly affects the cutting process. The working angles defined in the orthogonal plane can better reflect the cutting performance of the cutter. Taking the workpiece in Section 2.4 as an example, the cutter with a plane rake face is also designed as shown in Fig. 7, and the comparison with the curved rake face is shown in Fig. 8. For convenience, the cutter with curved rake face is named cutter A, and the cutter with plane rake face is named cutter B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002542_icem49940.2020.9270965-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002542_icem49940.2020.9270965-Figure3-1.png",
        "caption": "Fig. 3. Optimization of DWFT-PMM stator slot [5].",
        "texts": [
            " Meanwhile, to avoid the irreversible demagnetization in PM and burnout in power electronics, limiting the SC current below the safety threshold value is the straightest compensation measure. High reactance machine has inherent ability to limit the SC current in proper scope of electrical and thermal. The relation between stator parameters in PM machine and the slot leakage inductance, suppression effect on SC current, have been deduced [5]. In the 2D model, the narrow and deep slot type has higher slot leakage reactance which restrained the SC current, smaller b0 and bigger h0, showed in Fig. 3. The relation between SC current magnitude and the location of fault within the coil has been investigated by FEM and experimental tests [6, 7]. For stranded coils, the worstcase SC current generated when the SC fault happened close to slot opening. Adopting the vertically placed strip winding (VSW) could minimize the position dependency of the SC currents resulting from turn-turn fault and limit the SC current inherently. Meanwhile, low-rotor-pole-number machines have a better fault tolerance capability, while high-rotor-polenumber machines are lighter and provide higher efficiency"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003678_a:1008896010368-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003678_a:1008896010368-Figure4-1.png",
        "caption": "Figure 4. Simple model for determination of \u03bb.",
        "texts": [
            " (28) Using\u03bbu as the control input, \u03bbl and\u03bbr are controlled according to the following premise: Contact between the feet and ground is preserved without slippage when \u03bbl and \u03bbr satisfy the following conditions: |\u03bbr x | \u2264 \u00b5|\u03bbr y |, \u03bbr y \u2264 0, |\u03bblx | \u2264 \u00b5|\u03bbly |, \u03bbly \u2264 0, (29) where \u00b5 is the coefficient of friction between the ground and a foot. In order to determine the constrained forces under these conditions, the control input \u03bbu is determined so that \u03bbl and \u03bbr satisfy the following conditions. \u03bbr x = \u2212kr\u03bbr y, kr > 0, \u03bblx = \u2212kl\u03bbly, kl > 0, (30) where kr and kl must be smaller than the coefficient of friction \u00b5, so as to prevent slippage. Moreover, for the determination of kr and kl , we consider the simple model shown in Fig. 4, in which a mass m is suspended by wires. The coefficients kr and kl are determined to be the ratio of the tensile forces in the direction of the x and y axes. kr = L \u2212 x y , kl = x y . (31) According to Eq. (31), the length of the step, L , must be below a certain value in order to avoid slippage, because |kr | and |kl | must be smaller than \u00b5. Moreover, as the ground becomes slicker, L must be decreased. This tendency mimics that of a human taking short steps on a slippery surface. Although the model in Fig. 4 is a static model, it is sufficient for planning of the constrained forces at low walking speeds, (when the double support phase is critical.) When walking speed is high, the percent of time spent in the double support phase with respect to the whole walking cycle is low, and accurate control of the constrained forces is not required. Variables p\u0308r , p\u0307r , p\u0308l , and p\u0307l , contained in A and B, are ignored in the experiments shown in the next section, because their values are assumed to be sufficiently small during the double support phase"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000270_s12541-019-00298-4-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000270_s12541-019-00298-4-Figure2-1.png",
        "caption": "Fig. 2 Part design and dimensional information for experiments",
        "texts": [
            " The printing nozzle is then moved down to z-axis height of 0\u00a0mm; and the distance between nozzle and bed is adjusted tight enough to slide a standard printer paper. The wingnuts are adjusted so that the paper can barely move back and forth. The procedure is repeated for each corner, between corners, and the center of the bed. Third, the printing surface is covered with polymeric films and water-soluble glue. After the calibration is finalized, a test part is printed to check the first layer so that the adjacent layers stay even. Figure\u00a0 2 shows the part design used for the experiments, which is a dog bone shaped specimen with 9.00\u00a0cm (length) \u00d7 1.00\u00a0 cm (width) \u00d7 0.40\u00a0 cm (height). The dimensions of the specimen in the original study were 9.00\u00a0cm \u00d7 0.50\u00a0cm \u00d7 0.40\u00a0cm [60]; however, the width was modified to 1.00\u00a0cm to avoid the part breakage. It is noted that the printed part is not a commercial part and chosen solely to examine the capability of the FFF process for Fig. 1 Orientation angle of a printed part PLA, which aims to investigate whether FFF can produce sufficiently compliant replicates with suitable dimensional accuracy"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure13-1.png",
        "caption": "Fig. 13. Motion state diagram of 3-R(RRR)R+R HAM: (a) Pitch motion; (b) azimuth motion.",
        "texts": [
            " The length of members Pi Ji , Ji Gi ,Gi Ki and Ki Qi is L = 0.21m, and the radius of the fixed platform O and the moving platform C is R = 0.15m. Given the material of each component in ADAMS software, the masses and moments of inertias of 3-R(RRR)R+R HAM are measured as shown in Table II. According to the pitching and azimuth trajectory and rotation characteristics of the polarization mechanism shown in Table I, ADAMS software is used to carry out dynamic simulation of the HAM and the obtained pitching and azimuth motion states are shown in Fig. 13. According to the analysis in Section 1, the polarization mechanism needs to rotate and the polarization rotation driver needs to work in azimuth motion. The mass and inertia parameters of the mechanism in Table II are introduced into Eq. (37) to obtain the generalized inertia matrix of each member of the HAM under the initial pose. The pitch and azimuth motion parameters of the HAM described in Table I are transformed into the pose matrix of the moving platform, and the actuation data of each branch are obtained according to the inverse kinematics"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001776_tie.2020.3005108-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001776_tie.2020.3005108-Figure3-1.png",
        "caption": "Fig. 3. EMF vectors (electrical degree). (a)12s10r. (b)12s14r. The calculation method for the distribution factor kd of VFRM can be given as follow [26]:",
        "texts": [
            " The shadows on the DC field aligned with the rotor poles move with the rotor, i.e., these shadow areas are modulated by the rotor poles. The fundamental wave component of these modulated shadow areas is 2 pole-pairs, as shown by the dashed line. The modulated wave is 2 pole-pairs in space with 10 electrical cycles in every mechanical cycle. The 10 rotor poles modulating 6 pole-pairs of DC field result in the 2 polepairs of the rotary armature field. The coil-EMF vectors and phase coils for 12s10r and 12s14r VFRM are shown in Fig. 3. \ud835\udc58\ud835\udc51 = sin(\ud835\udc44\ud835\udc63\ud835\udf00/2) \ud835\udc44\ud835\udc60\ud835\udc56\ud835\udc5b(\ud835\udc63\ud835\udf00/2) (1) where Q is the number of least EMF vector per phase, \u03b5 is the angle between two adjacent vectors, and v is the order of harmonic. However, the armature windings are alternated wound on stator tooth, and its distribution factor is 1. For the non-modular VFRM, the single-coil back EMFs is the vector sum of two adjacent slot conductors, whose angular difference \u03b8c can be obtained for the vth back-EMF harmonic as \ud835\udf03\ud835\udc50 = 2\ud835\udf0b\ud835\udc63 | \ud835\udf0f\ud835\udc60 \ud835\udf0f\ud835\udc5f \u2212 1| (2) where \u03c4r=2\u03c0/Nr is the pole pitch, \u03c4s=2\u03c0/Ns is the slot pith"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002077_j.mechmachtheory.2020.104209-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002077_j.mechmachtheory.2020.104209-Figure5-1.png",
        "caption": "Fig. 5. Displacements of the Exechon PKM in x -direction: (a) Horizontal layout, (b) Vertical layout.",
        "texts": [
            " Table 3 Gravity-caused displacements of the Exechon PKM in x -direction (units: mm). Model type Horizontal layout Vertical layout Interval model ( \u03b5 I Px ) [-1.980, -1.897] [0.459, 0.469] Numerical model ( \u03b5 Px ) -1.927 0.461 Fig. 6. Interval displacements in x -direction of the Exechon PKM: (a) Interval method, (b) Scanning method The static analysis results of the PKM with nominal parameter values at two typical configurations of horizontal and vertical layouts (with \u03b8 = 0, \u03c8 = 0 and z = 1.35m) are illustrated as the follow. As shown in Fig. 5 , the direction of applied gravity ( g ) is set along x axis when the PKM is at the horizontal layout ( g = [1, 0, 0] T ), while at vertical layout the gravity is set along z axis ( g = [0, 0, 1] T ). The elastic displacements of the centre point P of the platform are extracted and analyzed only in x -direction. The simulation results obtained from the numerical model are compared with those computed by the proposed CIF based interval model as listed in Table 3 . As can be observed from Table 3 , the analytical results obtained from the interval model agree well with those of the numerical model in that the simulated elastic displacements in x -direction fall into the predicted intervals"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002765_s11071-021-06327-0-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002765_s11071-021-06327-0-Figure1-1.png",
        "caption": "Fig. 1 Side view of a tooth groove",
        "texts": [
            "1 Contact model There are three shapes of the sprocket tooth profile: convex, flat and concave. Compared with the previous two types, the concave profile has less compressive stress and fewer wear on the contact surface when interacting with the track under the same working condition. Therefore, it is widely used in tracked vehicles. According to the geometry of the tooth groove profile, it is assumed to consist of three cambered surfaces: the left, bottom and right. These three surfaces are numbered 1, 2 and 3 in turn. Figure 1 shows a side view of a groove, which is bilaterally symmetrical. Oj is the center of the sprocket, rp the radius of its dedendum circle and ra the radius of its addendum circle. O1, O2 and O3 are the centers of the left, bottom and right surfaces, respectively. Their corresponding radii are r1, r2, r3. Obviously, r1 \u00bc r3. The actual engagement between a single-pin track and a sprocket is illustrated in Fig. 2a. Contact forces can be assumed to occur between track pins and tooth grooves. According to different contact positions, there are three contact situations: the left surface contact, the bottom surface contact and the right surface contact, as exhibited in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002721_j.aej.2021.01.012-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002721_j.aej.2021.01.012-Figure7-1.png",
        "caption": "Fig. 7 Contact lines presented on worm hob surface with particular axial profile: a) rectilinear, b) concave, and c) convex.",
        "texts": [
            " The normal vector in the case of the rectilinear axial profile of the worm hob is described as n 2 0\u00f0 \u00de 1 \u00bc n 2 0\u00f0 \u00de x1 n 2 0\u00f0 \u00de y1 n 2 0\u00f0 \u00de z1 2 66664 3 77775 \u00bc L 2 0 1 0 \u00f0@r 1 0\u00f0 \u00de 1 @u1 @r 1 0\u00f0 \u00de 1 @u \u00de \u00f011\u00de where L 2 0 1 0 is the transformation matrix from system 1 0 to 2 0 , which is obtained by deleting the last line and the last column of the homogeneous matrix of equation (10), and this transformation matrix can be shown as L 2 0 1 0 \u00bc cos\u00f0u0 1\u00de sin\u00f0u0 1\u00de 0 cos\u00f0u0 2\u00de sin\u00f0u 0 1\u00de cos u 0 2 cos u 0 1 sin\u00f0u0 1\u00de sin u 0 1 sin\u00f0u0 2\u00de cos\u00f0u0 1\u00desin u 0 2 cos\u00f0u0 2\u00de 2 64 3 75 \u00f012\u00de The normal vector of the surface in the case of the arc axial profile of globoid worm is provided by n 2 0\u00f0 \u00de 1 \u00bc n 2 0\u00f0 \u00de x1 n 2 0\u00f0 \u00de y1 n 2 0\u00f0 \u00de z1 2 66664 3 77775 \u00bc L 2 0 1 0 \u00f0@r 1 0\u00f0 \u00de 1 @u1 @r 1 0\u00f0 \u00de 1 @h \u00de \u00f013\u00de The tangential vector is calculated based on the kinematics of worm wheel machining and is expressed as v 2 0\u00f0 \u00de 1 \u00bc v 2\u00f0 \u00de x1 v 2 0\u00f0 \u00de y1 v 2 0\u00f0 \u00de z1 2 6664 3 7775 \u00bc dr \u00f020 \u00de 1 du0 2 \u00bc dM 0 2 0 1 0 du0 2 r \u00f010 \u00de 1 \u00f014\u00de After solving equation (8), a set of solutions u1 for given values of parameter u (for the rectilinear profile) or a set of solutions u1 for given values of parameter h (in the case of the arc profile), is obtained. These parameters determine where the linear contact of the worm hob and worm wheel occurs [4]. If they are inputted into the tool surface equation, one gets the contact lines presented in the tool x 0 1y 0 1z 0 1 system. r \u00f010 \u00de c \u00bc r 1 0\u00f0 \u00de 1 u1; u\u00f0 \u00de \u00f0rectilinear profile\u00de r \u00f010 \u00de c \u00bc r 1 0\u00f0 \u00de 1 u1; h\u00f0 \u00de \u00f0arc profile\u00de \u00f015\u00de The contact lines for the tool of a particular profile are presented in Fig. 7. The generation of the worm wheel tooth flank surface is demonstrated in Fig. 8. In the considerations concerning the determination of the worm wheel mathematical model, the worm model is turned in relation to the z 0 1 axis so that the extreme cutting edge of the tool is located in the central plane, that is, plane y1z1 \u00bc y2z2 (Fig. 1). This position of the worm model is called the basic position. This operation simplifies the presentation of the mathematical model of the worm wheel tooth flank",
            " The surface generated during processing by the tool external edge and presented in the processing worm system is expressed by equation (19): r 1 0\u00f0 \u00de 2 \u00bc M 1 0 1 M12 M2 0 2 M21 r 1 0\u00f0 \u00de 1 u1\u00bcu1p\u00f0 \u00de \u00f019\u00de In matrix M 1 0 1 of equation (19), it is necessary to select the range of parameter u1 to obtain the surface of the worm wheel tooth flank of a given face width. This is done using the values of u1p and u1k. In Fig. 8b, the surface generated during processing by the tool external edge and determined based on equation (19) is presented. It is necessary to separate regions I and III from the surface shown in Fig. 8b. The boundaries of the zones are the contact lines lying in the area of the external edge of the cutting tool (Fig. 7) [19]. For region I, it is the contact line that does not lie in the tool axial plane (r \u00f010 \u00de c2 \u00bc r \u00f010 \u00de c i; 2\u00f0 \u00de) (Fig. 7). The separated zones I and III of the worm wheel tooth flank surface generated during processing by the tool external edge are illustrated in Fig. 8d. The algorithm for selecting region I (r 1 0\u00f0 \u00de 2I \u00de consists of checking the condition r 1 0\u00f0 \u00de 2x 1 0 i; j\u00f0 \u00de < r \u00f010 \u00de c1x10 i; 1\u00f0 \u00de \u00f020\u00de where r 1 0\u00f0 \u00de 2x 1 0 i; j\u00f0 \u00de is the element from the table of coordinates x 1 0 of the worm wheel side surface r 1 0\u00f0 \u00de 2 determined based on (19) and generated during machining by the tool extreme edge, r \u00f010 \u00de c1x10 i; 1\u00f0 \u00de is the element of the table of coordinates x 1 0 of the contact line r \u00f010 \u00de c1 (Fig. 7), and i; j are natural numbers. Equation (20) defines the range of coordinates i; j\u00f0 \u00de of the table for region I. Thus, it can be written as follows: r 1 0\u00f0 \u00de 2I i; j\u00f0 \u00de \u00bc r 1 0\u00f0 \u00de 2 i; j\u00f0 \u00de \u00f021\u00de where coordinates (i; j) meet the condition of equation (20). The algorithm for separating region III (r 1 0\u00f0 \u00de 2III ) is analogous. It consists of verifying the condition r 1 0\u00f0 \u00de 2x 1 0 i; j\u00f0 \u00de > r \u00f010 \u00de c2x10 i; 2\u00f0 \u00de \u00f022\u00de where r \u00f010 \u00de c2x10 i; 2\u00f0 \u00de is the element of the table of coordinates x 1 0 of the contact line r \u00f010 \u00de c2 (Fig. 7). Equation (22) specifies the range of coordinates i; j\u00f0 \u00de of the table for region III. Thus, it can be expressed as r 1 0\u00f0 \u00de 2III i; j\u00f0 \u00de \u00bc r 1 0\u00f0 \u00de 2 i; j\u00f0 \u00de \u00f023\u00de where coordinates (i; j) meet the condition of (22). Regions I, II, and III of the worm wheel flank surface have to be joined (Fig. 8e): r 1 0\u00f0 \u00de 2 \u00bc r 1 0\u00f0 \u00de 2I [ r 1 0\u00f0 \u00de 2II [ r 1 0\u00f0 \u00de 2III \u00f024\u00de The joining consists of assigning to vector r 1 0\u00f0 \u00de 2 i; j\u00f0 \u00de the elements of particular regions and reindexing the elements. The process of determining regions I, II, and III of the worm wheel generated by the tool of arc profile is analogous"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002066_01691864.2020.1854115-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002066_01691864.2020.1854115-Figure3-1.png",
        "caption": "Figure 3. The exoskeleton can support the wearer\u2019s forearm and follow the internal/external rotation of the wearer.",
        "texts": [
            " the forearm support, as shown in Figure 2(b). Themotor, pulley-timing belt and harmonic gear are aligned in parallel in order to reduce the thickness of the joint and the exoskeleton. Figure 2 presents the configuration of the active elbow joint. The flexion/extension movement of human elbow joint is sometimes accompanied with internal/external rotation of the shoulder joint. Thus, a passive revolute joint, Joint 5 with a telescopic mechanism (Figure 1 and 2(c)) of the exoskeleton, follows the wearer\u2019s movement as shown in Figure 3. Its length is adjustable, ranging from 0.50 to 0.65m (the length of l5 + l6 + l7, as shown in Figure 1(b) and Table 1), in order to fit the upper arms of different wearers. However, the passive revolute Joint 5 is not completely free. An appropriate rotational resistance is exerted to this joint to preserve the posture of the exoskeleton by tightening a screw on the outer part of the rotation shaft when the wearer performs a task that does not involve the internal/external rotation. In this way, the rotation center of the exoskeleton\u2019s elbow joint (Joint 6 in Figures 1 and 2) can be adjusted to coincide with the rotation center of the wearer\u2019s elbow joint to fit for the different wearers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002268_012047-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002268_012047-Figure4-1.png",
        "caption": "Figure 4. The finite element model of compressed airless tire.",
        "texts": [
            " The tread material property of airless tire model was investigated by compresion test according to ASTM D575 standard, while material property of spoke was investigated by tensile test according to ASTM D412 standard, respectively. The tread, wall and spoke properties were defined by Mooney-Rivlin hyperelastic model. On the other hand, the steel belt is specified to be an elastic isotropic material as described in table 2. The airless tire model was compressed with the compressive load of 1,000 kg. This simulation performs by moving the rigid flate plate to press airless tire model according to the testing of tire stiffness tester (Figure 4). The compressive load and vertical deformation of airless tire model was recorded while the airless tire model is compressed by rigid flat plate. The simulation result has a good agreement with the experimental result (EXP) from tire stiffness testing. The comparison of simulation model and experiment is presented in Figure 5. The result of using homogenization approach with isometric material different as the result of using REBAR element to replace the steel belt as shown in Figure 5. The vertical stiffness of airless tire finite element model by using homogenization approach and REBAR element were 905"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001371_s00170-019-04669-z-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001371_s00170-019-04669-z-Figure2-1.png",
        "caption": "Fig. 2 Height and width bead dimensions",
        "texts": [
            " After acquiring the data cloud, the three sets of beads resulted in 72 individual beads that were manually segmented. Each of those beads was then sectioned into a specific number of cross sections. Because the camera recorded images at 50 fps, there is a theoretical number of recorded frames for each bead which depends on its travel speed, as seen in Table 2. Those numbers of sections were then extracted from each bead with the use of the GOM Inspect software. From those sections, beadwidth and height weremeasured, according to their definition in Fig. 2. Bead height is here defined similarly to the literature, as the largest distance between substrate level and the bead\u2019s surface, normal to the substrate. The bead width dimension, however, is slightly different. It is defined as the largest distance parallel to the substrate between points on the clad bead surface that has at least 5% of the total clad bead height. This definition was incorporated on the measurement algorithm so that the extraction of those dimensions could be performed automatically and accurately"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003959_s0997-7538(99)00135-7-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003959_s0997-7538(99)00135-7-Figure1-1.png",
        "caption": "Figure 1. Idealised model of the tilted parametrically excited pendulum.",
        "texts": [
            " Theoretical justification to account for the stabilisation of this solution has been considered via an analytic treatment of the Mathieu equation (Phelps and Hunter, 1966) together with an approach based on an effective potential well (Michaelis, 1985). Here for the first time we extend the numerical investigation of the latter in considering the effect of introducing a tilt in the line of forcing to the system. In particular we identify the bifurcations which occur as forcing parameters are varied beyond the value for which initial stabilisation occurs. * Correspondence and reprints. The idealised model of the tilted parametrically excited pendulum is shown in figure 1 consisting of a point mass m suspended by a rigid light rod of length l. The pendulum is free to swing in the plane with the pivot point P subject to forced consinusiodal oscillations. We assume that the displacement of the pivot point, as measured from the fixed point O , describes a straight line offset by the angle \u03c6 from the upward vertical. Thus, at time \u03c4 , the pivot point has vertical coordinate z(\u03c4)= Z cos \u03c4 and horizontal coordinate x(\u03c4)=X cos \u03c4 . Including an overall simple linear viscous damping approximation, with damping constant \u03b6 , the equation of motion describing the planar motion of the pendulum is l d2\u03b8 d\u03c4 2 + \u03b6 d\u03b8 d\u03c4 \u2212 g sin \u03b8 \u2212Z 2sec\u03c6 sin(\u03b8 \u2212 \u03c6) cos \u03c4 = 0, (1) where \u03b8(\u03c4) is the angular displacement of the pendulum measured from the upward vertical position"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001752_j.mechmachtheory.2020.103992-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001752_j.mechmachtheory.2020.103992-Figure14-1.png",
        "caption": "Fig. 14. Contact stress distributions. (a) Concave side of the pinion, (b) convex side of the gear, (c) convex side of the pinion, and (d) concave side of the gear.",
        "texts": [
            " Stress analysis Based on finite element analysis, a loaded contact analysis can be completed to evaluate the stress distributions of the pinion and gear in Fig. 11 . The finite element model is as shown in Fig. 12 , which is meshed by the first-order hexahedral element. For the elements on the tooth surface, the max size of element length is set as 0.2 mm, and for other elements, the max size is set as 2 mm. The Young\u2019s modulus is 2.1 \u00d7 10 11 Pa and Poisson\u2019s ratio is 0.267. The torque of 540 Nm applied on the gear and the pinion is fixed as shown in Fig. 13 . The results of finite element analysis are shown in Fig. 14 . It can be seen that the shape and orientation of the contact ellipse of the finite element analysis are similar to those of the analytical analysis in Section 5.3 , which were used to verify the analytical model. Further, Fig. 14 also shows that the contact stress of the convex side of pinion is much larger than that of the concave side. This result has been predicted in Section 5.3 , because the length of the minor axis of the contact ellipse on the convex tooth surface of the pinion is generally less than that on the concave tooth surface of the pinion. 6. Conclusions In this article, to establish a systematic geometry design and analysis method of PCSBGs, an analytical synthesis about the kinematic geometry of these gears was completed based on the idea of tensor analysis and surface theory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001156_j.conengprac.2019.104118-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001156_j.conengprac.2019.104118-Figure3-1.png",
        "caption": "Fig. 3. Defined coordinate systems for the upper structure and equipment. The generalized coordinates \ud835\udf4d = [ \ud835\udf131 \ud835\udf132 ]\ud835\udc47 specify the configuration of the system. Additionally, arbitrary positions vectors which point exemplary to the IMUs and the position vectors of the different coordinate frames are depicted.",
        "texts": [
            " The position of a point P can be defined as a vector \ud835\udc58\ud835\udc91\ud835\udc58any \u2208 R3 with respect to the frame \ud835\udc58. The left superscript of the position vector denotes the frame, in which the vector is represented. The right superscript is the origin of the vector and the right subscript stands for the location of the point whose motion is of interest. For instance, the position vector 0\ud835\udc911P1 describes the position of the point P1 from the coordinate frame 1 resolved within the coordinate frame 0. Considering the material handling excavator as illustrated in Fig. 3, the frames are indexed as \ud835\udc58 = 0 for the upper structure, \ud835\udc58 = 1 for the boom and \ud835\udc58 = 2 for the stick. The position vector \ud835\udc58\ud835\udc91\ud835\udc58any can be transformed into the frame \ud835\udc57 using [ \ud835\udc57\ud835\udc91\ud835\udc57any 1 ] = \ud835\udc7b \ud835\udc57\ud835\udc58 [ \ud835\udc58\ud835\udc91\ud835\udc58any 1 ] , \ud835\udc57 \u2208 {0, 1, 2} (1) with the homogeneous transformation matrix (Siciliano et al., 2009, Sec. 2.7) \ud835\udc7b \ud835\udc57\ud835\udc58 = [ \ud835\udc79\ud835\udc57 \ud835\udc58 \ud835\udc57 \ud835\udc8d\ud835\udc57\ud835\udc58 \ud835\udfce 1 ] . (2) The rotation matrix \ud835\udc79\ud835\udc57 \ud835\udc58 \u2208 R3\u00d73 defines the rotation from the frame \ud835\udc58 to the frame \ud835\udc57 and the vector \ud835\udc57 \ud835\udc8d\ud835\udc57\ud835\udc58 \u2208 R3 describes the position of the frame \ud835\udc58 origin with respect to the frame \ud835\udc57 origin, resolved in the frame \ud835\udc57",
            " Applied to the material handling excavator, the homogeneous transformation matrices are formulated as \ud835\udc7b 0 1 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 cos(\ud835\udf131) \u2212 sin(\ud835\udf131) 0 0\ud835\udc5901,\ud835\udc65 0 0 \u22121 0\ud835\udc5901,\ud835\udc66 sin(\ud835\udf131) cos(\ud835\udf131) 0 0\ud835\udc5901,\ud835\udc67 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (3) and \ud835\udc7b 1 2 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 cos(\ud835\udf132) \u2212 sin(\ud835\udf132) 0 1\ud835\udc5912,\ud835\udc65 sin(\ud835\udf132) cos(\ud835\udf132) 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 , (4) where \ud835\udf131 \u2208 R describes the orientation from the upper structure to the boom and \ud835\udf132 \u2208 R the angle between the boom and the stick. The origins of the coordinate frames are set to coincide with the joints. Fig. 3 illustrates all used frames, parameters and joint angles. Any point 1\ud835\udc9111,any on the boom, expressed in the boom frame 1, can be transformed to the upper structure frame 0 using [ 0\ud835\udc9101,any 1 ] = \ud835\udc7b 0 1 [ 1\ud835\udc9111,any 1 ] . (5) A point 2\ud835\udc9122,any on the stick, represented in the stick frame 2, can also be transformed into the frame 0 with [ 0\ud835\udc9102,any 1 ] = \ud835\udc7b 0 2 [ 2\ud835\udc9122,any 1 ] (6) applying the matrix product \ud835\udc7b 0 2 = \ud835\udc7b 0 1\ud835\udc7b 1 2. The velocity or the acceleration of a point of interest are calculated with the first or second time derivative of the position vector 0\ud835\udc910any (Ballaire, 2015; Gross, Hauger, Schr\u00f6der, Wall, & Govindjee, 2014), i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002822_j.aime.2021.100040-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002822_j.aime.2021.100040-Figure12-1.png",
        "caption": "Fig. 12. Flow simulation using a baffle plate (red arrow) and a honeycomb grid (red circle), build plate position marked by red line.",
        "texts": [
            " Such flow behavior is insufficient for the PBF-LB/M process and would lead to impairments within the process zone. In a next step, a centrally mounted baffle plate was used with the aim of preventing flow separation and thus achieving a laminar fluid flow over the process zone. As shown in Fig. 11, flow separation also occurred, but in this case only in the later geometrical course of the flow. In addition, there was a reduction of the turbulent part in the flow. If a honeycomb grid with a depth of 25 mm, wall thickness of 0.1 mm and wall height of 3.2 mm (Fig. 12 (a)) was added to the model in the area of the main gas inlet (Fig. 12 (b)), the flow in the area of the process zone could be successfully kept laminar. For the design of the gas flow system, this model should therefore be implemented in the machine design. To compare the previous given designs, the fluid speed was extracted from the simulations, 2 mm above the area where the build platform is located. The speed distribution is shown in Fig. 13 and Table 3 provides the average speed and standard deviation of fluid speed. The simulation shows that the introduction of the baffle lowers the main flow of the gas and with this increases the average speed but degrades the homogeneity of the flow, measured by the standard deviation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001790_j.precisioneng.2020.06.014-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001790_j.precisioneng.2020.06.014-Figure6-1.png",
        "caption": "Fig. 6. Schematic representation of the reference coordinate system defined to perform CMM measurements of the micro-drilled holes. Circles measured for each cylinder at three different Z coordinates ( 1 mm, 2 mm and 3 mm) are displayed.",
        "texts": [
            " The micro-holes investigated in this work were measured using a tactile Coordinate Measuring Nachine (CMM) Zeiss Prismo Vast 7 (Zeiss, Oberkochen, Germany), with Maximum Permissible Error (MPE) for length measurements stated by the machine manufacturer equal to (2.2 \u00fe L/300) \u03bcm (where L is the measured length expressed in mm). The geometrical characteristics of interest were the holes diameter (nominally equal to 1.6 mm) and the cylindricity. The reference coordinate system was defined as illustrated in Fig. 6, with the Z-axis corresponding to the normal of the base plane (i.e. X\u2013Y plane), the x-axis corresponding to the line joining points a and b (defined as intersection between the cylinder axis and base plane), and the Y-axis oriented on the basis of the right-hand rule. A ruby spherical probe with diameter of 0.6 mm was chosen to measure three circles for each hole, within planes parallel to the base plane at three different Z coordinates: 1 mm, 2 mm and 3 mm (see Fig. 6). A total of 200 points were probed in scanning mode for each single circle, and the same points (i.e. 600 points coming from the three circles) were used to fit a cylinder to compute diameter and cylindricity. Five measurement repetitions were conducted to assess the measurement repeatability. The quality of the outer edges at the micro-holes entrance can be hindered by the presence of burrs remaining after the drilling process as well as other geometrical irregularities induced, for instance, by misalignment, vibration and run-out of the tool"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002666_s12369-020-00733-x-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002666_s12369-020-00733-x-Figure1-1.png",
        "caption": "Fig. 1 Handle location into the system",
        "texts": [
            " (1, 2), and applying the Laplace transform, the transfer functions that relate the linear and angular velocities with the applied force and torque can be obtained, f (S) = mSv(s) + bv(S) \u21d2 ( v(S) f (S) ) = ( 1 mS+b ) (3) \u03c4(S) = i S\u03b8\u0307 (S) + c\u03b8\u0307 (S) \u21d2 ( \u03b8\u0307 (S) \u03c4 (S) ) = ( 1 i S+c ) (4) With these transfer functions, the robot\u2019s end effector Cartesian velocity can be obtained as a result of the force and torques applied by the user, \u23a1 \u23a3 vx vy vz \u23a4 \u23a6 = ( 1 mS+b ) \u23a1 \u23a3 fx fy fz \u23a4 \u23a6 , \u23a1 \u23a3 \u03b8\u0307x \u03b8\u0307y \u03b8\u0307z \u23a4 \u23a6 = ( 1 I S+c ) \u23a1 \u23a3 \u03c4x \u03c4y \u03c4z \u23a4 \u23a6 (5) The forces and torques used in the previous equations are in the robot\u2019s end effector and relative to the robot\u2019s base frame. In the use case described in Sect. 3 where the hand guidance has been implemented and tested it is not the case (Fig. 1). Thehandle to use the handguiding interface is not attached directly to the robot\u2019s end effector. It is attached to a mechanical device that allows the use of the tools and the hand guiding interface at the same time (Fig. 1). The orientations of the handle are not coincident with the robot\u2019s end effector orientations. Moreover, the forces and torques measured on the sensor are on its own reference frame. For these reasons, the forces and torques measured on the F/T sen- sor ( \u0302( f f t ), \u0302(T f t )) have to be transformed to the robot\u2019s end effector and relative to the robot\u2019s base frame (\u0302( f R), \u0302(T R)) . This can be done using the rotation matrix from robot\u2019s base frame to end effector frame R Ree , that can be obtained from the homogeneous transformation matrix calculated with the forward kinematics of the robotic arm, and using the rotation matrix that describes the orientation of the F/T sensor relative to the end effector ee R f t , obtaining the respective forces: \u23a1 \u23a3 fx ee fyee fzee \u23a4 \u23a6 = (ee R f t ) \u23a1 \u23a3 fx f t fy f t fz f t \u23a4 \u23a6 \u21d2 \u23a1 \u23a3 fx R fy R fz R \u23a4 \u23a6 = (R Ree) \u23a1 \u23a3 fx ee fyee fzee \u23a4 \u23a6 (6) where fx ee is the force expressed in the robot\u2019s end effector frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003134_s11012-021-01388-2-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003134_s11012-021-01388-2-Figure2-1.png",
        "caption": "Fig. 2 Equivalent mechanism for the waterbomb pattern described by a unit cell showed in light blue",
        "texts": [
            " 1c) and a cylindrical origami (Fig. 1d). The following sections present a local asymmetry study, evaluating the behavior of a unit cell of the waterbomb pattern considering both kinematics and mechanical approaches. The analysis of the waterbomb pattern can be based on the hypothesis of rigid origami, where all the deformation is localized on the creases (mountain and valley folds), which means that faces remain flat and undeformed. Therefore, it is possible to analyze the waterbomb unit cell as a mechanism (Fig. 2), where the creases are represented by revolute (or cylindrical) joints and the faces are represented by rigid links. Waterbomb pattern is a spherical mechanism where its movement is restricted to a sphere. Based on that, it can be described as a 6R linkage mechanism [3, 5] with mobility 3 [9, 10]), meaning that 3 variables are necessary to fully describe the position of the mechanism. The kinematics of a unit cell can be described using this rigid 6R mechanism, allowing to define the number of inputs required to completely describe the unit cell, treated as a linkage mechanism",
            " Therefore, the transformation matrix from joint i and to joint i\u00fe 1 is given by iT i\u00fe1 \u00bc cos hi\u00f0 \u00de cos hi\u00f0 \u00de sin hi\u00f0 \u00de cos ai\u00f0 \u00de sin hi\u00f0 \u00de sin ai\u00f0 \u00de ai cos hi\u00f0 \u00de sin hi\u00f0 \u00de cos hi\u00f0 \u00de cos ai\u00f0 \u00de cos hi\u00f0 \u00de sin ai\u00f0 \u00de ai sin hi\u00f0 \u00de 0 sin ai\u00f0 \u00de sin ai\u00f0 \u00de cos ai\u00f0 \u00de Ri 0 0 0 1 2 664 3 775 \u00f02\u00de where ai is the angular distance between two consecutive joints, from zi to zi\u00fe1 axis about the xi\u00fe1 axis; hi is the rotation of the ith joint, from xi to xi\u00fe1 axis about the zi axis; ai is the offset distance measured from the origin Oi to the intersection of axes zi and xi\u00fe1, along the xi\u00fe1 axis; and Ri is the joint offset, measured as the distance from the i frame to the intersection of axes zi and xi\u00fe1, along the zi axis (Fig. 3). The waterbomb pattern has a characteristic that all joints intercept at a common point (point O in Fig. 2), resulting in ai \u00bc Ri \u00bc 0 \u00f0i \u00bc 1. . .6\u00de. In addition, ai is fixed for each pair of consecutive joints, being associated with the angle k that defines the shape of the waterbomb cell and, for a squared waterbomb cell, k \u00bc p=4. The D\u2013H formulation allows the description of each i ! i\u00fe 1 joint pair through just four parameters and since three of them are constant values, each joint can be represented by one degree of freedom, hi, resulting into 6 free variables. Since the waterbomb pattern is related to a closed chain, the last joint is connected to the first one",
            " The analysis is performed considering a closed m nwaterbomb tessellation, withm lines and n cells on each line (Fig. 14). The tessellation is bFig. 11 Workspace for cases P1 and P2 and simulations for a generic case (asymmetric), plane-symmetric behavior (P1, P2) and symmetric behavior (P3). The correspondent deformation of each origami face is shown for cases P3 (b), P2 (c), P1 (d) and asymmetric (e) symmetrically actuated by considering that the cells are pulled radially through the middle vertex of each unitary cell (point O in Fig. 2). Longitudinal symmetry is placed on the unit cells vertex, in such a way that the radius of each line measured from the tessellation axis to the vertex (point O in Fig. 2) is the same between lines and within the same line. This analysis is performed in two stages: the first study evaluates the influence of the number of lines (m) considering a tessellation with n = 6 cells on each line; the second study evaluates the influence of the number of cells on each line (n) considering two tessellations: with m = 5 lines and with m = 6 lines. In order to evaluate the influence of the number of lines (m), four tessellations formed by waterbomb unit cells on each line are of concern varying the numbers of lines (Fig",
            " Acknowledgements The authors would like to acknowledge the support of the Brazilian Research Agencies CNPq, CAPES and FAPERJ. Declaration Conflict of interest The authors declare that they have no conflict of interest. Appendix This appendix presents details about the origami formulation. Equivalent mechanism analysis The waterbomb unit cell is a closed-loop mechanism and, for the formulation used in this paper, it is assumed that the first linkage i \u00bc 1\u00f0 \u00de is associated with the crease OB, being numbered counterclockwise. Therefore, the last linkage i \u00bc 6\u00f0 \u00de is related to the crease OA (see Fig. 2). The frame definition is summarized as: 1. The first frame i \u00bc 1\u00f0 \u00de is defined as the creaseOB. 2. Frames are disposed following a counterclockwise sequence, following the vertex order B; C; D; E; F and A, starting from i \u00bc 1\u00f0 \u00de at vertex B and ending at i \u00bc 6\u00f0 \u00de at vertex A. 3. The zi axis of each i frame is aligned with the crease, with the origin at O (see Fig. 2). 4. The yi axis of each frame is coplanar with the origami face delimited by the joints i and i 1, and the y1 frame is coplanar with the origami face delimited by frames 1 and 6. 5. The xi axis of each i frame is the normal of the face delimited by joints i and i 1, and the x1 frame is the normal to the face delimited by frames 1 and 6. 6. The waterbomb defines an inner region and an outer region, where the inner region is contained within the waterbomb edges AB; BC; CD; DE; EF and FA. Each xi axis points outwards the inner region",
            " With this consideration, zi axis associated with valley folds (creases OA; OC; OD andOF) are positioned along the crease, pointing from Oi to the correspondent vertex (A; C; D or F), while zi axis associated with mountain folds (OB and OE) are positioned along the crease, pointing to the opposite direction of the correspondent vertex (B or E). The values of the D\u2013H parameters for a generic waterbomb cell are given in Table 3. The waterbomb pattern has a characteristic that all joints intercept at a common point (point O in Fig. 2), resulting in ai \u00bc Ri \u00bc 0 \u00f0i \u00bc 1; . . .6\u00de. In addition, ai is fixed for each pair of consecutive joints, being associated with the angle k that defines the shape of the waterbomb cell wherein, for a squared waterbomb cell, k \u00bc p=4. Finite element analysis The behavior of origami structures is described assuming quasi-static equilibrium, where the shape change is due to a succession of equilibrium configurations. It is assumed that the total potential energyU is the sum of the strain energy stored in bars, Ubar , the strain energy stored in folding (torsional springs on the creases) and bending (torsional springs as virtual folds), Uspr, the work done by external loads, Vext, By considering quasi-static equilibrium, the ith bar element is represented by Ti bar \u00bc AiLiSx oEx oui Ki bar \u00bc AiLi Sx o2Ex ou2i \u00fe C oEx oui oEx oui T \" # \u00f09\u00de where Sx is the second Piola\u2013Kirchhoff (P\u2013K) tensor, C is a tangent modulus, Ai is the transversal section area of the bar element and Li is the length of the bar element"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001485_j.mechmachtheory.2019.103776-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001485_j.mechmachtheory.2019.103776-Figure3-1.png",
        "caption": "Fig. 3. D-H parameters of an infinitesimally mobile RCRCR linkage [43] . The red arrows x 1 , y 1 , z 1 denote the reference coordinate systems, other blue arrows represent the body-fixed coordinate systems of links. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " (31) for this linkage also only involves one self-stress coefficient \u03b2 as Q = \u03b2 ( 2 a a a 0 ) , (48) which is sign indefinite with two eigenvalues of different sign. Thus, the corresponding linkage possesses higher-order infinitesimal mobility, and is actually a finite mechanism. The two eigenvalues represent two bifurcation branches in the current singular configuration in Fig. 2 (b). It is known that the same results could be obtained based on a matrix method in point coordinates [5 , 41] . 6.2. Infinitesimal mobile RCRCR linkage The shaky property of this RCRCR linkage in Fig. 3 was reported by [42] using input-output relationship of joint angles and reinvestigated by [43] using acceleration analysis based on screw theory. The D-H parameters of this linkage are shown in Fig. 3 and below, \u03b112 = 3 \u03c0/ 4 , a 12 = 1 , \u03b82 = \u03c0/ 2 , d 2 = 2 \u221a 2 ; \u03b132 = \u03c0/ 2 , a 23 = 2 \u221a 2 , \u03b83 = 7 \u03c0/ 12 , d 3 = 1 + \u221a 3 ; \u03b134 = \u03c0/ 2 , a 34 = 0 , \u03b84 = \u03c0/ 2 , d 4 = 3 ; \u03b154 = \u03c0/ 6 , a 45 = 2 / \u221a 3 , \u03b85 = 0 , d 5 = 8 \u2212 3 \u221a 3 ; \u03b151 = 5 \u03c0/ 6 , a 51 = 1 / \u221a 3 , \u03b81 = 0 , d 1 = 4 \u221a 3 \u2212 3 . (49) The quantity \u03b1ij and a ij in Eq. (49) represents the twist angle and the link length between the axes of joint i and joint j , respectively, \u03b8 i represents the joint angle about the joint axis i , and d i represents the axial offset along joint axis i . The direction of twist angles and joint angles are shown in Fig. 3 . All these parameter are defined according to D-H notation [44] . The joints are indexed with the numbers of the body fixed-coordinate systems. The screw coordinates of joints in the given configuration are \u03be11 = ( 0 0 1 0 0 0 )T , \u03be12 = ( 0 0 0 0 0 1 )T , \u03be2 = ( 0 \u2212 \u221a 2 / 2 \u2212 \u221a 2 / 2 0 \u221a 2 / 2 \u2212 \u221a 2 / 2 )T , \u03be31 = ( \u22121 0 0 0 0 \u22124 )T , \u03be32 = ( 0 0 0 \u22121 0 0 )T , \u03be4 = ( 0 \u221a 3 / 2 \u22121 / 2 2 \u2212\u221a 3 / 2 \u22123 / 2 )T , \u03be5 = ( 0 1 / 2 \u2212\u221a 3 / 2 2 \u221a 3 \u2212 3 / 2 \u22121 / 2 \u2212\u221a 3 / 6 )T , (50) in which \u03be11 , \u03be12 are the screw coordinates of the cylindrical joint 1, representing one revolute joint and one prismatic joint respectively, and \u03be31 , \u03be32 are the screw coordinates of the cylindrical joint 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001019_rpj-07-2018-0182-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001019_rpj-07-2018-0182-Figure11-1.png",
        "caption": "Figure 11 Mapped contact pressure plot in static analysis",
        "texts": [
            " In the second step (folding), rotational and translation degrees of freedom is arrested for both horizontal axis and the top die is translated vertically downward till it closes completely by touching bottom die and pressure applied at the inner surface of the tube is held constant. Field mapping option available in FEA software is used to transfer the contact pressure distribution from the explicit analysis to a static analysis where only the bottom die is considered for analysis, while other boundary conditions remain same. Contact pressure obtained during the forming is considered as a static load and applied to die in the static analysis. Figure 11 shows the mapped surfaces, boundary condition and load because ofmapping. Contact stress distribution plot for the explicit analysis is shown in Figure 12. Maximum contact stress is observed where the Metal bellow hydroforming Prithvirajan R. et al. Rapid Prototyping Journal D ow nl oa de d by N ot tin gh am T re nt U ni ve rs ity A t 0 2: 47 3 1 M ay 2 01 9 (P T ) bellow root region will be formed and reduces towards the crest forming region and lesser at other regions where there is no contact"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002903_j.engfailanal.2021.105453-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002903_j.engfailanal.2021.105453-Figure2-1.png",
        "caption": "Fig. 2. The location and structure of rubber joint (At the right is a cross section of rubber joint).",
        "texts": [
            " Once the components of bogie have problems, it will have an affect on the safe operation of rail vehicle and may result in huge loss of life and property. Therefore, it is very important to analyze the failure of rail vehicle and propose a reasonable solution. The rubber universal spherical joint (hereinafter referred to as rubber joint) studied in this paper is an important suspension component of the bogies of one subway (Fig. 1). The schematic diagram of its location and structure are shown in Fig. 2 and the related data are shown in Table 1. The rubber joint is composed of axle, rubber bush and steel cover. The rubber material makes the rubber joint have a good shock absorption function, while the inner and outer metal materials provide sufficient strength and stiffness. The subway is composed of 6 sections and its maximum running speed is 80 km/h. The operating temperature of the rubber joints is \u2212 15 \u25e6C ~ +40 \u25e6C. The designed service life of the rubber joints is 5 years. However, in<2 years, the rubber parts were found to be sticky and cracked during a routine maintenance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000183_tim.2019.2949319-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000183_tim.2019.2949319-Figure5-1.png",
        "caption": "Fig. 5. Pillar layout and channel tapering in the linear coupler DCW TWT. Inset: assembled structure with its WR-28 waveguide flanges and hole for electron beam injection/extraction.",
        "texts": [
            ", in the longitudinal direction of the TWT, is easier to fabricate, but requires a longer structure with the consequent need for a larger volume magnetic field. The second, novel, approach is based on bending the DCW through 90\u00b0. This permits the positioning of tapered coupling sections orthogonally to the interaction section. The manufacture and S-parameter characterization of these two approaches to the DCW TWT are described below. The linear coupler DCW SWS topology consists of an uniform interaction section with both ends terminated with sections where the pillar height is linearly tapering to zero (Fig. 5). This enables wideband coupling between the TE10 mode of rectangular waveguide and the hybrid mode supported by the DCW. The optimized DCW comprises a 40 period interaction region and two linear DCW transitions connected by a bent waveguide to the WR-28 flanges. Each DCW transition includes 15 pillar rows linear tapered in height. The waveguide enclosure has a sinusoidal tapered profile toward the bend to improve the matching. The structure is fabricated in two parts. One is the DCW with the couplers as shown in Fig. 5, and the other is a plane metal lid. Hex-head machine screws are used to fix the two parts together. Aluminum was chosen as metal for the first prototype for its easier machining in comparison with copper. The structure was machined on a KERN Pyramid milling machine. A photograph of part of the interaction region is shown in Fig. 6. The S-parameters of the realized DCW were measured and compared with the 3-D electromagnetic simulations. Good agreement, Fig. 7, is obtained for S11 and S21. In particular, the maximum frequency difference is of about 1%"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure14-1.png",
        "caption": "Fig. 14. Journal bearing eccentricity under load conditions (a) coordinate system; (b) definition of mobility vector.",
        "texts": [
            " The load applied to journal bearings is supported by the lubrication films between the journal and the bearing. In this paper, Mobility Method (modified from [35,36]) is adopted to model the journal bearing response under dynamic loadings. Based upon this model, the trajectory of journal motion inside the bearing can be integrated by the integrator of the dynamic system. This method uses a graphical method \u2018mobility\u2019 to analyze the journal center orbit marching in time from some initial eccentricity ratio on the mobility maps, as illustrated in Fig. 14. The mobility vector is defined on a \u03be-\u03b7 coordinate, in which \u03be-axis is aligned with the direction of the applied force. Given the current position of the journal \u03be and \u03b7, the loading force vector, the dimension of bearing and clearance, and viscosity of the lubricants, the instantaneous velocity can be determined as \u03be\u0307 = F ( 2Cr / Djb )2 \u03bcLjbDjb W\u03be \u2212 \u03b7\u03c9 \u03b7\u0307 = F ( 2Cr / Djb )2 \u03bcLjbDjb W\u03b7 + \u03be\u03c9 (45) where the components of mobility vector W\u2192 can be found graphically from the mobility map. For typical EGPs, the short-journal bearing model is more applicable than the long-bearing one"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002592_iros45743.2020.9340957-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002592_iros45743.2020.9340957-Figure1-1.png",
        "caption": "Fig. 1. A group of robots jointly manipulating an object along a desired trajectory. The robots compute a collision free trajectory, even if only a subset of the robots know about the presence of obstacles through local sensing.",
        "texts": [],
        "surrounding_texts": [
            "I. INTRODUCTION\nIn many situations, manipulating an object requires multiple robots. We present the Scalable Optimal Collaborative Manipulation with Local Constraints (SOCM LoCo) algorithm, a distributed algorithm by which a group of robots can collaboratively move an object to a desired configuration while avoiding obstacles in the environment. Each robot only communicates with its neighbors over a connected communication network, making the algorithm distributed. The algorithm\u2019s efficient scalability arises from its constant computational complexity, independent of the number of robots, allowing for an arbitrary number of robots to collaborate. The algorithm also allows for individual robots to sense obstacles online and impose collision avoidance constraints without other robots explicitly knowing about these constraints. In simulation, we show that SOCM LoCo provides 100 times better tracking performance and several orders of magnitude faster convergence rate than a distributed method based on Consensus alternating direction method of multipliers (ADMM) and performs comparably to leaderfollower methods.\nOur algorithm would be useful in a variety of collaborative manipulation applications, including in automated manufacturing or warehouse environments and in autonomous construction of buildings or structures in hazardous or remote environments such as in space. In these applications, a team of robots can work together to move large parts or sub-assemblies into place, providing greater scalability than traditional monolithic robots. In addition, our algorithm can\n1,2 O. Shorinwa and M. Schwager are with the Department of Mechanical Engineering and Department of Aeronautics and Astronautics, Stanford University, USA.\nThis work was funded in part by DARPA YFA grant D18AP00064, NSF NRI grant 1830402, and ONR grant N00014-18-1-2830. We are grateful for this support.\nbe applied in disaster relief scenarios, where heavy and varied debris must be cleared by a group of robots to aid emergency workers in search of survivors. In all these cases, our algorithm enables the coordination of the group of robots in manipulating objects.\nThe algorithm proceeds in a receding horizon fashion. Each robot solves a series of local optimization problems iteratively and communicates with its neighbors over a communication network, enabling the entire group to manipulate the object along an optimal trajectory respecting the object\u2019s dynamics, collision avoidance constraints, and input constraints of the robots. For convex constraints and affine object dynamics, our algorithm is guaranteed to converge to the globally optimal trajectory. In more practical cases with non-convex constraints such as collision avoidance and non-linear dynamics, including rotational dynamics, the algorithm gives locally optimal, collision free trajectories. The algorithm is inspired by the Separable Optimization Variable ADMM (SOVA) method [1]. Specifically, we use a separable optimization variable property unique to the collaborative manipulation problem to greatly reduce the number of optimization variables handled by each robot, thereby leading to faster convergence and better performance than other ADMM methods.\nThe contributions of this work are as follows. We derive the SOCM LoCo algorithm and prove its convergence to the globally optimal trajectory for an object with affine dynamics and convex constraints. Through iterative linearization, we extend the method to the case of non-linear dynamics and non-convex constraints. Further, we show the algorithm has a constant computational complexity, independent of the number of robots, while other methods scale cubically in the number of robots. In simulation, we show SOCM LoCo converges many orders of magnitude faster and attains a tracking error 100 times smaller than distributed methods based on Consensus ADMM.\n978-1-7281-6212-6/20/$31.00 \u00a92020 IEEE 9108\n20 20\nIE EE\n/R SJ\nIn te\nrn at\nio na\nl C on\nfe re\nnc e\non In\nte lli\nge nt\nR ob\not s a\nnd S\nys te\nm s (\nIR O\nS) |\n97 8-\n1- 72\n81 -6\n21 2-\n6/ 20\n/$ 31\n.0 0\n\u00a9 20\n20 IE\nEE |\nD O\nI: 10\n.1 10\n9/ IR\nO S4\n57 43\n.2 02\nAuthorized licensed use limited to: Rutgers University. Downloaded on May 17,2021 at 19:56:37 UTC from IEEE Xplore. Restrictions apply.",
            "This paper is organized as follows: in Section II, we note previous approaches and formulate the manipulation task as an optimization problem in Section III. We derive our distributed algorithm using the alternating direction method of multipliers in Section IV and apply our algorithm to objects with non-convex dynamics in Section V. In Section VI, we demonstrate our algorithm with as many as 100 robots collaboratively manipulating an object under different communication constraints. We provide concluding remarks in Section VII.\nMany approaches achieve collaborative manipulation using centralized control schemes which aggregate local information from each robot to compute individual motor commands [2], [3], [4]. These approaches require significant computation and communication and do not scale with the number of robots involved in the manipulation task. To allow for distributed schemes, previous approaches use potential fields to guide each robot to the object and subsequently manipulate the object by enclosing it, described as object closure or caging [5], [6]. These methods do not provide a mechanism for specifying local constraints for each robot which our method provides. To handle constraints, some methods plan for collision free trajectories in convex spaces of the environment centrally and follow these trajectories using local controllers [7], [8]. Our method does not require any centralized computation procedure, making it fully distributed.\nSome other approaches employ a leader-follower architecture in which a single robot (leader) actively manipulates the object along a desired trajectory while the other robots (followers) infer the motion of the leader and move in consistency with the leader\u2019s motion. In some approaches [9], the followers communicate their motion constraints explicitly to the leader which shares the desired trajectory to all robots, incurring significant communication and poor scalability, while others require a human operator to issue commands for the group of robots [10]. Some other methods allow for collaboration between the leader and followers without communication. In these approaches, only the leader knows its desired trajectory. The followers estimate the leader\u2019s trajectory using force sensors and move along with the leader using non-linear feedback controllers [11], [12], [13], impedance controllers [14], [15], [16], and adaptive controllers [17]. For small groups of robots, the followers can estimate the leader\u2019s trajectory using impedance controllers without force sensors [18], with greater tracking errors for larger groups. All the above approaches allow the robots to follow the leader\u2019s trajectory but do not address trajectory planning for the group of robots. In contrast, our method finds the optimal trajectory for the object, avoiding collisions in the environment.\nOther distributed methods describe the desired object trajectory using impedance without a designated leader [19]. The desired object impedance is distributed among the robots to compute individual control inputs using knowledge of the\ngeometric relations between the robots. Each robot receives its desired trajectory before the task and follows the specified trajectory through impedance control. Like leader-follower approaches, these methods allow for following a desired trajectory without trajectory planning. In addition, these methods are suited to manipulation tasks involving a few robots (2 to 5 robots) with known grasp points and do not scale to large groups of robots.\nOur method enables a group of robot to collaboratively plan an optimal trajectory for an object, responding to avoid obstacles in the environment known by a subset of the robots, while manipulating the object. In addition, with constant computational complexity, our algorithm scales to large groups of robots. We derive our algorithm using the alternating direction method of multipliers (ADMM), exploiting separability of the optimization variable to achieve lower computation and communication complexity. ADMM has been applied in receding horizon control, albeit in a centralized fashion [20]. Our distributed method provides greater tolerance to errors through feedback with a receding horizon approach.\nWe desire to manipulate an object from an initial configuration to a desired final configuration by controlling a group of robots to work in collaboration. Each robot grasps the object before manipulating it. We denote the object\u2019s configuration and velocities as xobj which includes its position and translational velocity and its orientation and angular velocity. We consider a group of N robots manipulating the object. Robot i applies force Fi and torque \u0393i to the object. We can express the manipulation task as an optimization problem given by\nminimize \u222b T\nt=0\n( \u03c8(xobj(t)) + N\u2211 i=1 \u03b2i(Fi(t),\u0393i(t)) ) dt\nsubject to g(xobj(t), F (t),\u0393(t)) = 0 \u2200t h(xobj(t), F (t),\u0393(t)) \u2264 0 \u2200t\n(1) We include a desired trajectory tracking objective for the manipulated object \u03c8(\u00b7) and encode desired behaviors for robot i such as to minimize energy or fuel consumption in \u03b2i(\u00b7). In addition, we consider dynamic constraints and equality constraints on the initial configurations and velocities of the object and robots given by g(\u00b7). We also incorporate additional convex constraints on the object\u2019s configuration and its derivatives denoted by h(\u00b7) which can include collision avoidance constraints represented by safe convex zones within the environment.\nCommunication Graph\nWe represent the robots as nodes in an undirected graph G described by a set of vertices V = {i | i = 1, \u00b7 \u00b7 \u00b7 , N} and a set of edges E . An edge (i, j) exists in E if robot i can communicate with robot j. We assume the communication graph G is connected i.e. we can obtain a path linking every pair of nodes from the edges in E . In addition, we denote the\n9109\nAuthorized licensed use limited to: Rutgers University. Downloaded on May 17,2021 at 19:56:37 UTC from IEEE Xplore. Restrictions apply."
        ]
    },
    {
        "image_filename": "designv11_14_0000276_s00366-019-00893-z-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000276_s00366-019-00893-z-Figure11-1.png",
        "caption": "Fig. 11 Perspective view of the four-module tensegrity structure (a), the basic quadruplex module (b) and top view of the structure",
        "texts": [
            "\u00a09) for varying values of the friction coefficient. The tension values shown correspond to a 5\u00a0kN concentrated load applied at each of the middle nodes of the structure. Increasing the friction coefficient alters internal force distribution in the four sub-elements of the sliding cable. The maximum deviation from the friction-free tension value approaches 7%. In this example, the folding of a four-module tensegrity tower is simulated. The structure is composed of four identical quadruplex modules with four bar-to-bar connections (Fig.\u00a011). The structure has 16 struts held in space by 36 cables. The vertical cable elements running through the four modules are grouped into four sliding cables that control the configuration of the structure. Perspective and top views of the tensegrity structure are presented in Fig.\u00a011a, c where dashed lines denote the four continuous cables. Figure\u00a011b illustrates the basic quadruplex module. The tensegrity structure has a nominal height of 200\u00a0cm. Struts have a length of 102.5\u00a0cm and are made of aluminum hollow tubes. All struts have a cross-section area of 2.5\u00a0cm2 and a Young\u2019s Modulus of 7500\u00a0kN/cm2. For folding purposes, the middle saddle cables at \u00bc and \u00be height levels are replaced by spring elements with a constant of 2\u00a0kN/cm and a free length of 54\u00a0cm. Top, middle and bottom saddle cables have a rest length of 80\u00a0cm. The four sliding cables have a rest length of 256\u00a0cm",
            " Cable-length adjustments are deliberately performed in small steps to eliminate dynamic effects. Figure\u00a012 illustrates three snapshots of the folding process of the structure. The folding of the tensegrity structure is analyzed using the two LC-enhanced methods. Sample results obtained for four actuation steps are compared. The values of the height 1 3 of the tensegrity structure as well as the tension forces in two cables (cable 13 and cable 37) are presented in Table\u00a07. Cable 13 connects node 1 to node 5 (see Fig.\u00a011) and cable 37 connects node 17 to node 13 (see Fig.\u00a011). Both cables are sub-elements belonging to the same sliding cable. Results are displayed for the LC-enhanced transient stiffness method and the dynamic relaxation method considering a frictional coefficient of 0.1. Once again, it is noticed that the two LC-enhanced methods yield similar results for the folding of the tensegrity structure. The differences between the results of the two methods are less than 1% for all tested configurations. As for the computational efficiency, each actuation step employing the LC-enhanced dynamic relaxation needs 0",
            " Figure\u00a013 illustrates the evolution of the tension in element 13 with respect to the actuation step during folding of the structure. Sliding-induced friction causes the tension in cable 13 to increase progressively during folding process. For a coefficient of friction equal to 0.4, the internal force in cable 13 is approximately 10% larger than the reference friction-free predicted value. Figure\u00a014 illustrates the evolution of the tension force in element 37 during the folding of the structure. Element 37 is the sliding cable sub-element connecting node 17 to node 13 (Fig.\u00a011). Recall that element 37 and element 13 are the two extreme sub-elements of the same sliding cable. Sliding-induced friction causes the tension force in cable 37 to decrease progressively during the folding process compared with the reference friction-free predicted values. This is more pronounced at fully folded configuration for the structure. With a friction coefficient that is equal to 0.4, the internal force in cable 37 is approximately 8% lower than the reference friction-free predicted value"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003936_s0076-6879(98)96047-5-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003936_s0076-6879(98)96047-5-Figure1-1.png",
        "caption": "FIG. 1. Solution paths toward and away from the rotating disk electrode (RDE). (A) Vertical cross section of the R D E is shown with arrows indicating the flow of solution from a field perpendicular to the electrode surface and exit of solution from the electrode in a field parallel to the plane of the electrode. (B) Radial diffusion paths (due to centrifugal force) across the electrode surface, shown in an end-on view of the electrode. (C) Diffusion path occurring within the RDE/ incuba t ion chamber used in the studies of transporter functioning.",
        "texts": [
            " diffusion coefficient of the electroactive species in cm2/sec, C is the concentration of the electroactive species in the bulk solution in mol/liter, u is the kinematic viscosity of the solution in cm2/sec, ~0 is the angular velocity of rotation in rad/sec, where ~o = 2rrN, and N is the number of rotations per second. The RDE theory is based on considering an infinitely thin plane, the electrode, rotating in solution at a constant rate about its axis. The motion of the liquid, caused by the drag (stirring) of the electrode, pulls the solution to the electrode surface from a field perpendicular to the electrode. The analyte, dissolved in solution, is brought to the electrode and then spun radially away horizontally by centrifugal force (Fig. 1). If a potential sufficient to oxidize (or reduce) the dissolved analyte is applied to the electrode, then electron flow to (or from) the electrode is detected as current. This electrolysis at the electrode surface drives the concentration of the analyte to zero, producing a thin diffusion layer. Diffusion across this layer is driven by the concentration gradient between the solution at the electrode surface and the bulk solution, and therefore must remain constant since the concentration gradient across the diffusion layer is constant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure21-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure21-1.png",
        "caption": "Fig. 21. Semi-Flexible Type Traction Coupling",
        "texts": [
            " In many such cases the coupling is also required as a load-limiting device, so that the driven shaft can be stalled while the motor continues to run at, say, one-and-a-half to twice full-load torque for a short period, until the cause of the overload is removed, or an overload trip with a delay characteristic cuts off the power. It will be recognized that a sufficient drag torque is essential, as the means by which the motor sets the driven shaft in motion. The two most sat is factory and simple methods of reducing drag torque are those employed 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 111 in the semi-flexible type traction coupling, Fig. 21, namely the reservoir chamber on the back of the runner and the anti-drag baffle near the inner profile diameter. When the runner is stalled the working circuit partially empties into the reservoir, thus reducing the torque, and further, the baffle interferes with the circulatory flow of the remaining liquid so that the drag torque is reduced to about one-quarter that of the original fluid flywheel or plain Vulcan coupling. Slip Curves. The results are shown by the slip curves plotted at full torque on the basis of engine speed in ",
            "comDownloaded from PROBLEMS OF FLUID COUPLINGS 129 In other cases it is possible to adopt the \u201cflexible type\u201d of traction coupling which has no internal journal bearings and thus permits of both parallel and angular misalignment within the limits of the running clearances. In this case the engine crankshaft carries the weight of the impeller and casing, while the runner is overhung from the shaft of the driven machine to which it is bolted, by a rigid coupling. An intermediate solution is to use the \u201csemi-flexible\u201d type of traction coupling, Fig. 21, a description which is used where a flexible traction coupling with tie-rod thrust embodies a self-aligning ball bearing to support the runner shaft in the boss of the impeller. An ordinary outboard bearing can then be used with a flexible coupling on the driven shaft ; alternatively the outboard bearing is not required if the flexible coupling is of the self-centring type, or has a centring spigot so that the short runner shaft can act in effect like a cardan shaft and permit of considerable errors in alignment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000841_asjc.1990-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000841_asjc.1990-Figure1-1.png",
        "caption": "FIGURE 1 The block diagram of compensator based commmand filter",
        "texts": [
            " [37] The command filter is described as follows: z1 = \ud835\udf14z2, z2 = \u22122\ud835\udf17\ud835\udf14z2 \u2212 \ud835\udf14(z1 \u2212 \ud835\udefc1), (8) where \ud835\udf14 > 0 and \ud835\udf17 \u2208 (0, 1]. If the input signal \ud835\udefc1 fulfills |?\u0307?1| \u2264 \ud835\udf0c1 and |?\u0308?1| \u2264 \ud835\udf0c2 for all t \u2208 [0, +\u221e), where \ud835\udf0c1 > 0, \ud835\udf0c2 > 0 and z1(0) = \ud835\udefc1(0), z2(0) = 0, for any \ud835\udf07 > 0, there exist \ud835\udf14 > 0 and 0 < \ud835\udf17 \u2264 1, so we have |z1 \u2212 \ud835\udefc1| \u2264 \ud835\udf07, and |z\u03071|, |z\u03081| and|z...1| are bounded. While, each command filter is designed to computer z1 and z\u03071 without differentiation. The block diagram of command filter is shown in Figure 1. Remark 4. According to the characteristics of the first-order filter, the difference in finite time of the deviation signal approximates the differential of the deviation signal, resulting in enlargement of the signal noise. With the increase of the system order, the overall system will be instable. In order to deal with the filtering errors zi \u2212 \ud835\udefci, employing the error compensation mechanism at each step of the filters is necessary. Lemma 3. [43] If there have a continuous radially unbounded function V \u2236 Rn \u2192 R+ \u222a {0} such that: (i) V(x) = 0 \u21d4 x = 0; (ii) the solution x(t) satisfied the inequality V(x) \u2264 \u2212 (\ud835\udefcVp(x) + \ud835\udefdVq(x))k for some \ud835\udefc, \ud835\udefd, p, q, k > 0 \u2236 pk < 1, and qk > 1; then, the globally fixed-time stable can be achieved and the settling time T satisfies that T(x0) \u2264 1 \ud835\udefck(1\u2212pk) + 1 \ud835\udefdk(qk\u22121) ,\u2200x0 \u2208 Rn"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002161_9783527813872-Figure6.20-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002161_9783527813872-Figure6.20-1.png",
        "caption": "Figure 6.20 Schematic illustration of organic field-effect transistor (OFET) with the bottom gate, top contacts configuration; the applied polarization of electrodes corresponds to p-type semiconductor; L and W are the length and width of the channel, respectively, Z is the thickness of the gate insulator, and x is the space coordinate.",
        "texts": [
            " Such a strong dependence of the current on the applied voltage was found by Helfrich and Mark in thin single crystals of diphenyl, p-terphenyl, p-quaterphenyl, naphthalene, and anthracene [114]. The current density\u2013voltage response in the case of exponential distribution of traps energy is given by jSCLCt = N0\ud835\udf07e1\u2212l ( \ud835\udf00r\ud835\udf000l Nt(l + 1) )l(2l + 1 l + 1 )l+1 Ul+1 L2l+1 (6.29) where N0 is the density of energetic states and l = Tc/T > 1 (Tc is the characteristic temperature of the distribution energy of traps). In an electronic device schematically illustrated in Figure 6.20, consisting of a semiconducting layer with two electrodes (called source and drain) and a third electrode (called gate) isolated from the semiconductor, it is possible to fill energetic traps in the semiconductor by applying adequate voltages to the electrodes. If the states in the DOS tail (see Figure 6.17b) are filled, then the charge transport in the channel will be facilitated. This effect is used in OFETs, in which filling of traps is controlled by the gate potential. It is worth mentioning that a concept of field-effect transistor was patented already in the twenties of the 20th century [115]. In fact, Figure 6.20 shows a typical structure of OFET, with the bottom gate and top contact configuration. In OFET, the surface density of current flowing between the source and drain electrodes (jDS), if the voltage between these electrodes (UDS) is constant, depends on the density of charge carriers in the transistor channel. Taking into account equation for capacitor, one can easily show that the charge dq(x) accumulated in some area W dx (where W is channel width), close to the semiconductor/dielectric interface, at the distance x from the source electrode is equal: dq(x) = dC(x)(UGS \u2212V (x)), where dC(x) = Cdx/L is a capacity of a capacitor with infinitely small surface area W dx and C is the capacity of the capacitor formed by the gate electrode, dielectric, and charge layer induced in the transistor channel on area WL"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001297_0954405419883052-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001297_0954405419883052-Figure1-1.png",
        "caption": "Figure 1. Diagram of Gaussian Sphere.",
        "texts": [
            " For complex hybrid manufacturing process, the TAR calculation method should evaluate the dynamic five-axis movement capability of machine tool as well as the possible combinations of multiple alternative printing and machining. Definition of tool accessibility based on Gaussian Sphere For any particular tool contact point Pi on the part surface, in order to describe the tool accessibility, a unit hemisphere called a Gaussian Sphere can be attached at this point. Each point on the Gaussian Sphere represents one tool axis direction at Pi. Meanwhile, there is a specific continuous area on the sphere indicating the TAR at that point and the other area indicating the tool non-accessibility range (NTAR), as shown in Figure 1. For the additive process, the region is called printing accessibility range (PAR), which is used to characterize the accessibility of the printing tool. Similarly, for the subtractive process, the region is called cutting accessibility range (CAR), which is used to characterize the accessibility of cutting tool. Printing accessibility. Taking printability into consideration, the whole printing process should be free of collision and free of requirements for auxiliary support. Therefore, the PAR at a certain point can be expressed as the intersection of the non-collision range (NCR) and the self-supporting accessibility range (SAR) as follows PAR(Pi)=NCR(Pi)\\SAR(Pi) \u00f01\u00de In the additive process, the self-supporting printing can be realized by limiting the printing direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002851_09544089211011011-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002851_09544089211011011-Figure1-1.png",
        "caption": "Figure 1. Geometry and dimensions (in mm) of the sample part.",
        "texts": [
            " Maraging steel MS1 metal powder material is produced by the EOS manufacturer company. EOS MaragingSteel MS1 is a pre-alloyed ultra high strength steel in fine powder form. This kind of steel is characterized by having excellent mechanical properties and being easily heat-treatable using a simple thermal age-hardening process to obtain superior hardness and strength. Design and manufacturing of the sample part for dimensional accuracy The geometry and dimensions of the sample part manufactured by the DMLS device are given in Figure 1. The sample part is manufactured by layers manufacturing process with 0.04 mm constant layer thickness and Standard EOS manufacturing parameters. Images of the layers are captured by UI154xSEM model number camera with 1280x1024 px resolution, color depth of 32 bpp and exposure time of 78.03 ms. After the manufacturing is finished, the part is first cleaned from inside the manufacturing cabinet. Then, the support structures are removed after cutting the piece from the wire erosion device (Figure 2(a) and (b))"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001934_iros45743.2020.9341409-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001934_iros45743.2020.9341409-Figure6-1.png",
        "caption": "Fig. 6: Force body diagram for the Axel Rover. Tether-dependent static stability of the system depends on terrain contact points with the wheels (P1,P2), potential contact with the boom and the ground (P3), and the tether forces on the end of the boom (F3), and the tether tension applied. Correct static stability requires a knowledge of the position of the last anchor point.",
        "texts": [
            " This is determined by examining the position of the vector between X1 and aN in the coordinate frame, Fd, whose axes include the surface normal of the contact point of aN (red arrow) and the vector between aN and aN\u22121 (blue arrow). If the tether-anchor vector is above the XY plane of this frame, it is considered a detachment event. In the illustration, the state X0 does not contain a detachment while the state X1 does. Once the in-contact pose, N points of surface contact, and the state\u2019s anchor history have been determined, the pose stability is evaluated using constrained quadratic programming. This consists of searching for feasible contact and tether forces that place Axel in static equilibrium (Fig. 6). While the settler treats ground contact as a point intersection, the actual contact generally consists of a finite-sized contact patch enclosed by a pair of grousers. This allows each contact \u201cpoint\u201d to carry moments between Axel and the ground. This property is fundamental to Axel\u2019s ability to perform certain maneuvers such as lifting the boom off of the ground. Thus, 7037 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on June 19,2021 at 16:46:11 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000134_j.jfranklin.2019.07.020-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000134_j.jfranklin.2019.07.020-Figure1-1.png",
        "caption": "Fig. 1. Illustration of the coordinates in the Serret\u2013Frenet frame.",
        "texts": [
            " The second contribution is that a FTD has been employed to estimate the value f the lumped disturbance and compensate the controller in real time, which can significantly urther improve the robustness property of the closed-loop path-following system. The paper is organized as follows. Section 2 gives preliminaries and problem formulation. n Section 3 , the procedure of the composite finite-time controller is developed. Simulation tudies are conducted in Section 4 . Section 5 concludes this paper. . Problem formulation The Serret-Frenet frame, as shown in Fig. 1 , is widely adopted to study the straight-line ath-following problem of the USVs [33] . The origin of the frame { SF } is located at the closest point on the curve C from the origin f frame { B }. C is the given desired path, e is defined as the distance between the origin of SF } and { B }, \u03c8 is the heading angle of the ship and \u03c8 SF is the path tangential direction. he kinematics of path-following errors based on the Serret\u2013Frenet frame can be described s follows [34] \u02d9 e = u sin \u03c8 e + v cos \u03c8 e \u02d9 \u03c8 e = \u02d9 \u03c8 \u2212 \u02d9 \u03c8 SF = \u03ba 1 \u2212\u03bae ( u sin \u03c8 e + v cos \u03c8 e ) + r (1) Please cite this article as: Q"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001489_s40430-019-2151-7-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001489_s40430-019-2151-7-Figure1-1.png",
        "caption": "Fig. 1 Cantilever CNTRC trapezoidal plate",
        "texts": [
            " It is due to the mathematical complexities involved in geometry of trapezoidal plates and difficulties involved in free edges. So, in this paper, free vibration analysis of cantilever FG-CNTRC trapezoidal plates are studied. The plate is modeled based on FSDT, and the set of governing equations and boundary conditions are mapped from the trapezoidal area into a rectangular one. GDQM is employed as a numerical approaches, and natural frequencies and corresponding mode shapes are reported for various cases. As depicted in Fig.\u00a01, an FG-CNTRC cantilever trapezoidal plate clamped at y = 0 and free at other edges is considered. The plate is of dimensions a and b and angles \u03b1 and \u03b2 and is reinforced by CNTs arranged in y direction. As Fig.\u00a02 shows, four standard patterns of distribution of CNTs are considered including UD, FG-V, FG-O and FG-X. For these types of distribution, volume fraction of CNTs is given as [35] where h is thickness of the plate and V\u2217 CNT is total volume fraction of CNTs which is same in all types of distribution"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002489_j.jwpe.2020.101671-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002489_j.jwpe.2020.101671-Figure3-1.png",
        "caption": "Fig. 3. Part of the analysis carried out on this model before 3D printing. (1) An isometric figure of the anode chamber. (2) Top view of the anode chamber with the representation of the substrate flow. (3) Isometric view showing the point of entry (bottom) and exit (top) of the substrate.",
        "texts": [
            " Under the simulations, it was seen that this caused a negative effect by considerably lowering the speed of the fluid in the vicinity of the electrodes. For model (8), the effect of the disturbances was used with certain changes so that the fluid did not pass through the middle of the disturbance, but was forced to travel at a higher speed near the electrodes, instead. For the last model (9), the number of perturbations along the length and width of the anodic chamber was increased, thus increasing the speed of the substrate in the desired areas. Fig. 3 shows part of the analysis carried out on this model before 3D J.D. Lo\u0301pez-Hincapie\u0301 et al. Journal of Water Process Engineering 38 (2020) 101671 printing. Different steps for the construction of the selected MFC are shown in Fig. 4. A stock feed solution was prepared by mixing sodium acetate anhydrous (J.T.Baker, Fisher Scientific, Pennsylvania, USA) 132.98 g L\u2212 1; NH4Cl, 8 g L\u2212 1; MgCl2, 2 g L\u2212 1; CaCl2, 1 g L\u2212 1; KH2PO4, 1.2 g L\u2212 1; CuSO4 0.02 g L\u2212 1; ZnCl2, 0.02 g L\u2212 1; MnSO4, 0.02 g L\u2212 1. The theoretical Chemical Oxygen Demand (COD) of this mixture was 100 g L\u2212 1 (1 gCOD per 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001217_j.mechmachtheory.2019.103633-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001217_j.mechmachtheory.2019.103633-Figure3-1.png",
        "caption": "Fig. 3. ZSFH_SCM.",
        "texts": [
            " Furth more, IORFH has approximately linear stiffness characteristics. The inner and outer rings are rigid. So IORFH has excellent properties. It is used as the positive stiffness mechanism in this paper. Bi et al. [35] introduce a pre-pressured spring between the inner and outer rings of IORFH, that is the zero stiffness flexure hinge based on the spring-crack mechanism (ZSFH_SCM). One end of the spring is fixed on the inner ring (the point B i , i = 1 , 2 , 3 ) and the other is fixed on the outer ring (the point A i , i = 1 , 2 , 3 ), as is shown in Fig. 3 . The intersection point O of the three beam flexures is the center of the rings. The inner ring is equivalent to a crank, that is the dotted line O B i ( i = 1 , 2 , 3 ) . In other words, the crank length equal to the length of OB i . Thus, a negative stiffness mechanism is constructed. That is the spring-crank mechanism, in which the spring is designed as the diamond leaf spring string [35] . The diamond leaf spring string is made of n diamond leaf springs in series, as is shown in Fig. 4 . It is a compressible spring based on the beam deformation theory of Awta [38] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001759_j.mechmachtheory.2020.103995-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001759_j.mechmachtheory.2020.103995-Figure11-1.png",
        "caption": "Fig. 11. Trajectories of points on the kirigami form. (a) is front view; (b) is left view; (c) is top view.",
        "texts": [
            " For the kirigami form corresponding to the double crank, the angle \u03d5 increases all the time, but it will go through the bifurcation point where the panels p 1 and p 6 were coplanar with \u03d5= 180 \u00b0. This can be interpreted as kirigami form of type I can be directly bifurcated into type II through the bifurcation points B 3 and B\u2019 3 in Fig. 7 . The characteristics of the linkage are inherited by its kirigami form. Here, of particular interests are crank-rocker linkage and its kirigami form. The trajectories on the linkage are monitored as shown in Fig. 10 , while the trajectories on the kirigami form are plotted in Fig. 11 during the whole cycle of motion when panel p 1 is fixed. The subscript i of M i or N i indicates that the point is on the revolute joint i or the crease z i . Points located on the creases i in the kirigami form have similar trajectories to the points on the revolute joint i in the linkage. For example, M 5 and N 5 are different points on the revolute joint 5 of the linkage and crease z 5 of the kirigami form, respectively. Both the trajectories of M 5 and N 5 are circular arcs despite different starting point and arc length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001798_s00419-020-01733-z-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001798_s00419-020-01733-z-Figure5-1.png",
        "caption": "Fig. 5 Section view of housing, rotating shaft, and bearings",
        "texts": [
            " Bearing in the electric motor side in all cases of the test is healthy, but the bearings on the other side of the test bench can be healthy or defective during research. Bearings are installed in rigid housings that have the capability of applying axial preload. In Figs. 5 and 6, the schematic of the designed test bench, and also a section view of housing, a rotating shaft and bearings are shown. The two ends of the shaft are threaded to fix the positions of the inner race on the shaft by nuts and also to fix the position of the outer race on the housing by tightening the socket cap screws (please see Fig. 5). The electric motor and the rotating system are connected by a flexible coupling. The type of this coupling is \u201ccurved jaw flexible coupling\u201d which can absorb vibration, parallel, angular misalignments, and shaft end-play as shown in Fig. 6. The special grease KluberQuiet\u00ae BQ 72-72 which produces a considerably low-level noise is used for bearing lubrication. Before data collection, the environmental conditions are controlled and the necessary equipment is provided. The test bench is leveled to avoid applying axial loads, and the bearings and electromotor\u2019s axis are aligned with high-precision laser alignment equipment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003171_s42114-021-00327-9-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003171_s42114-021-00327-9-Figure3-1.png",
        "caption": "Fig. 3 CAD model of porous structures. (a, b) Tetrahedron model. (c, d) Diamond model",
        "texts": [
            " The metal powder will be collected in a sealed container located under the cyclone separator. The morphology of the metal powder is shown in Fig.\u00a02. The powder was spherical or nearly spherical, with a relatively smooth surface and good flowability. The particle size ranged from 12 to 53\u00a0\u03bcm. In this study, tetrahedral and diamond unit cells were generated using 3D modeling software. The tetrahedral and diamond unit cells were arranged linearly in the horizontal and vertical directions to construct porous structures with dimensions of 10 \u00d7 10 \u00d7 10 mm3, as shown in Fig.\u00a03. The porosity and strut size are shown in Table\u00a01. The porous structure was formed using SLM machine (SLM-250, independently developed by the North University of China) equipped with a Yb-fiber laser 400\u00a0W. The forming process is to spread the metal powder onto the processing substrate through the spreading device, and the laser will be controlled by the computer. When irradiated to the designated area, the metal powder in the irradiated area will melt, and the molten metal will quickly cool and solidify"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002542_icem49940.2020.9270965-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002542_icem49940.2020.9270965-Figure11-1.png",
        "caption": "Fig. 11. Flux paths in the HESM [31].",
        "texts": [
            " Experiment has shown a good magnetic isolation between sub-motors, so the failure of any sub-motor should not significantly affect the operation of others. 2) Flux Controllable PM Machine Besides the conventional rotor-PM machines, doublesalient PM (DSPM) machine, memory machine and flux switching PM (FSPM, in Fig. 10) machines have been investigated[29, 30]. These stator-PM machines have additional benefit, which PMs mounted in stator are easier to dissipate heat, especially in post-fault operation. Meanwhile, the hybrid excitation synchronous machines (HESM) for SG application are also get much attention recently [31, 32], in Fig. 11. The common thing in these types is controllable flux to amend in the presence of fault. The fault winding, especially short circuit, of the machine could benefit from the controllable flux. A doubly-salient flux controllable out-rotor motor topology has been presented in [27], Fig. 12. The innovation of this topology is based on double-salient motor and memory motor. Adopted the AlNiCo PMs can be completely demagnetized in the post-fault mode and the motor will operate as switched reluctance motor, which Authorized licensed use limited to: University of Gothenburg"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000234_ecce.2019.8913191-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000234_ecce.2019.8913191-Figure1-1.png",
        "caption": "Fig. 1. Exploded view of the proposed VFMM.",
        "texts": [
            " Moreover, LCF PMs are unintentionally demagnetized by applying maximum load current because of their low coercive force. The torque density is reduced by unintentional demagnetization. In addition, iron loss is tend to be increased when magnetization reversal is caused by negative magnetizing current pulse because the harmonic flux becomes larger. In this paper, a novel VFMM structure which improves abovementioned challenging points is proposed and simulated by two-dimensional finite element analysis (2D-FEA). The proposed VFMM has delta-type PM arrangement and the large flux barrier as shown in Fig. 1. Basic parameters of the proposed VFMM are similar to IPMSM mounted in TOYOTA PRIUS fourth-generation that is the commerciallysupplied hybrid vehicle in 2015 [16]. Incidentally, detailed characteristics of IPMSM mounted in TOYOTA PRIUS are measured and analyzed by our research team. 978-1-7281-0395-2/19/$31.00 \u00a92019 IEEE 6054 II. CHARACTERISTICS AND OPERATING PRINCIPLE OF VFMM VFMMs can be classified into DC or AC-magnetized motor by magnetization method. In general, DC-magnetized VFMM have winding to apply magnetizing current apart from the armature winding [17, 18]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001564_j.robot.2020.103482-FigureA.28-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001564_j.robot.2020.103482-FigureA.28-1.png",
        "caption": "Fig. A.28. On the left \u2013 schematic of a conventional wheel assembly, on the right \u2013 trace of the caster wheel contact point with respect to the robot base.",
        "texts": [
            " This indicates the system is holonomic for omnidirectional driving unless a desired support polygon velocity is imposed (bxcp = bxcp,des) together with the camber angle (\u03d5\u0307 = \u03d5\u0307des). In such situations, (A.5) can be expressed as (A.3) with a constant offset. Thus (A.5) is subject to the same non-holonomic constraint, when both the support polygon and camber angles are tracked. In this section, (33) is applied to analyse the evolution of the support polygon for well-studied conventional wheel assemblies: fixed, steering and caster wheels. To this end, the kinematics of the generalised standard wheel assembly, as given at Fig. A.28, has been considered first; (33) results to ox\u0307(s)cp|ti = [(R + r)\u03bd\u0307 0 0 ] . (B.1) In the next step, the feasible values of \u03bd\u0307 are computed from the NSPR constraint for each wheel type and applied to (B.1). Finally, to achieve more clear-cut result, the robot world posture is eliminated from (B.1) by moving the vector origin to the robot base with a rigid-body transformation that reads bx\u0307(b)cp = bRs (ox\u0307cp \u2212 ox\u0307b )(s) \u2212 ( \u03c9b \u00d7 oxb )(b) , (B.2) where oxb represents the current base position, \u03c9b \u2208 \u211c 3 denotes the angular velocity of the robot base with respect to the inertial frame, and bRs \u2208 \u211c 3\u00d73 expresses the rotation matrix from the steering frame to the base reference frame. Table B.5 provides a summary of the above analysis. It shows that the contact point placement is locked for the fixed and steering wheels, while the mobility of the castor wheel imposes the support polygon adjustment on a circular path around the wheel steering axis whose radius equals to the wheel offset (Fig. A.28). As follows from (33) the holonomic omnidirectional driving scheme can be designed for the wheeled platform provided that its support polygon is not constrained and/or the camber angle is not imposed. In case of standard wheeled platforms, that applies only to the systems with only castor wheels, whose motion imposes support polygon change. Other standard platforms, with fixed and steering wheels, enforce a constant support polygon where the former excludes omnidirectional driving, and the later introduces non-holonomy [13]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002591_scems48876.2020.9352321-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002591_scems48876.2020.9352321-Figure1-1.png",
        "caption": "Fig. 1. Framework of digital twin.",
        "texts": [
            " But an accepted view is that the DTT is a bridge between the physical world and the digital virtual world. Researchers believe that the DT is a digital model of a physical entity that is created digitally, using data to simulate the behavior of a physical entity in a real environment. Through virtual and actual interactive feedback, data fusion analysis, decision-making iterative optimization and other means to add or expand the new ability for physical entities [10]. A typical digital twin architecture is shown in Fig. 1. It can be found that DT mainly includes three parts: digital simulation model, physical entity, and information interaction module. The physical entity includes the interconnection and perception standard interface between the physical elements and supports plug and play. It has wide-area sensors and state feedback points, which can collect information with high density and wide frequency. Besides, the physical entity can also receive the optimization instructions from the data interaction module or the simulation data from the digital simulation model to change the physical element combination mode, production process, and resource matching"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure1.10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure1.10-1.png",
        "caption": "Fig. 1.10 Housing for UAVs: a housing cover; b automatic gate; c charging terminals; d UAV; e snowdrift",
        "texts": [
            " In this approach, electrical connection between ground and on-board circuits appears right after landing when UAV\u2019s undercarriage open electrodes touch and simply lay on ground platform open electrodes (pads, strips). Such solutions provide arrangement of charging process even in conditions of inaccurate UAV landing and charging for a group of UAVs simultaneously. Of course, open contact pads are affected by atmospheric precipitation and dust contamination. So conception of open contact pads usage includes also application of hangars for UAV service stations. Any hangar or other housing (Fig. 1.10) is necessary not only for contact pads protecting but for providing comfort temperature for Li-Po accumulator cells charging. It is well-known that most of LiPo batteries may be charged only under positive values of temperature. Therefore, covered temperature-stabilized housing is required for arranging of UAV service stations in winter conditions. Besides temperature stabilizing subsystem, such a housing must contain control subsystem for automatic opening/closing a gate in time of a next UAV takeoff/approaching"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001265_s12555-019-0101-x-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001265_s12555-019-0101-x-Figure2-1.png",
        "caption": "Fig. 2. Quadrotor with coordinate frames.",
        "texts": [
            " If we define a reference frame with the target in the center, the motion of the aircraft can be described by a standoff distance and a phase angle with respect to the target. Hence, it is more convenient to describe the standoff tracking motion by using cylindrical coordinate system, such as \u03c1\u0308 = a\u03c1 +\u03c1\u03c6\u03072, \u03c1\u03c6\u0308 = a\u03c6 \u22122\u03c1\u0307\u03c6\u0307, z\u0308 = az, (2) where \u03c1 is the standoff distance, \u03c6 is the azimuth angle, a\u03c1 , a\u03c6 , az are the radical, tangent and vertical accelerations, respectively. 2.2. Quadrotor dynamics The schematic of the quadrotor is presented in Fig. 2, which is detailed in [24]. The coordinates Xe, Ye, Ze denote the Cartesian inertial frame relative to the earth with its origin at a known point. The Cartesian coordinates Xb, Yb, Zb are fixed to the body and its origin coincides with the center of gravity of the vehicle. The motion of the quadrotor has six degrees of freedom which can be described by the position vector (x,y,z) and the Euler angle vector (\u03d5 ,\u03b8 ,\u03c8). The nonlinear dynamic model of the quadrotor is expressed by the following equations: mx\u0308 =U1(cos\u03c8 sin\u03b8 cos\u03d5 + sin\u03c8 sin\u03d5), my\u0308 =U1(sin\u03c8 sin\u03b8 cos\u03d5 \u2212 cos\u03c8 sin\u03d5), mz\u0308 =U1 cos\u03b8 cos\u03d5 \u2212mg, Ix\u03d5\u0308 = \u03b8\u0307 \u03c8\u0307(Iy \u2212 Iz)+ lU2, Iy\u03b8\u0308 = \u03d5\u0307 \u03c8\u0307(Iz \u2212 Ix)+ lU3, Iz\u03c8\u0308 = \u03d5\u0307 \u03b8\u0307(Ix \u2212 Iy)+ cU4, (3) where x, y, z are the translational positions in the Earthfixed frame"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure1.4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure1.4-1.png",
        "caption": "Fig. 1.4 Symbiotic UAV-UGV robotic team (Illustration: [14])",
        "texts": [
            " ARs of the same type are commonly used in groups, but some examples are known when different types of vehicles are included into the system [13]. Such solution allows to complement functions and possibilities of ARs (Fig. 1.3). Moreover, there are symbiotic aerial ground robotic teams where UAVs are used for aerial-specific tasks, while unmanned ground vehicles (UGVs) aid and assist them [14, 15]. UGVs can provide UAVs with a safe landing area, while UAVs can provide an additional degree of freedom for the UGVs, helping them to avoid obstacles (Fig. 1.4). Such a system may be used, for example, for parcel transportation scenarios. Number of ARs in the system may be huge. In 2018, a Chinese drone company has launched into the sky 1374 drones above the city of Xian to set a GuinnessWorld Record for most UAVs flown at the same time [16]. The quadcopters simultaneously took off and created various colorful 3D formations in the air (Fig. 1.5). Control Station. The control station (CS) for UAVs may be based on the ground (GCS), aboard ship (SCS) andpossibly in a parent aircraft (ACS)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003715_s0301-679x(98)00106-6-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003715_s0301-679x(98)00106-6-Figure7-1.png",
        "caption": "Fig. 7. Scheme of the reflection photoelastic measurements.",
        "texts": [
            " The sealing in the chamber is ensured on one side by the contact of the seal segment against the moving plate and on the opposite side by a properly sealing element. The housing, where the seal is mounted, has geometry and sizes according to constructor indications. Under the test chamber the moving plate is visible. During the tests a fluid pressure p=4 bar has been used. The photoelastic measurements have been carried out by means of a photoelastic polariscope with a white light source, a polariser and an analyser. The test scheme is represented in Fig. 7. The optic group polariser\u2013analyser can be rotated in a synchronous way by a precision mechanical transmission in order to show the isoclinic lines. A couple of quarter wave plates can be oriented to have a circular polariscope and to show the isochromatic fringes. The seal segment and the thin photoelastic sheet are schematically represented; the photoelastic material is fastened to the seal section. The photoelastic sheet is made of special material for measurements on rubber; it is the PS-6 Sheet and the glue is the PC-9 Adhesive of Measurement Group"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure14-1.png",
        "caption": "Fig. 14. Quick-Emptying Coupling with Centrifugal Transfer to Gravity Tank",
        "texts": [
            " In either of the above cases, the application of automatic control is simply a matter of regulating the valves in question, and the coupling is independent of auxiliary power. The rate of filling is, however, rather slow when the coupling is nearly empty, and a reduced quantity is then circulating through the scoop tube. Temperature effects are a factor to be reckoned with in the operation of both the suction scoop tube and the ejector method of filling, hence further schemes for regulating the coupling without auxiliaries were studied. Gravity Filling and Rim Pumping. One method, illustrated by Fig. 14, involved reversion to the old practice of gravity filling, the scoop tube being dispensed with in this case and provision made for very rapid emptying. The fluid coupling was arranged in a casing, of which the bottom part formed a shallow sump and the upper part a gravity tank immediately above the coupling. Quick-emptying valves were provided, of the kind shown by Fig. 20, which are held shut by the centrifugal head of liquid when the coupling is running. They are 7 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003126_s42835-021-00852-z-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003126_s42835-021-00852-z-Figure2-1.png",
        "caption": "Fig. 2 Structure of the air-core stator (the composite cylinder is not shown)",
        "texts": [
            " The thermal and mechanical of the one pole pair stator was analyzed by the finite element analysis (FEA). A pole pair full-scale oil-cooling air-core stator was designed, manufactured, and tested. A 2\u00a0MW DD HTS generator is mainly composed of the oil-cooling air-core stator, HTS magnet, torque transfer system, coolant transfer coupling, stator cooling system, and cryogenic refrigeration system, as shown in Fig.\u00a01. The oilcooling air-core stator is consisted of back iron, GFRP teeth, windings, and a composite cylinder, as shown in Fig.\u00a02. Especially, the composite cylinder locates in the inner radial surface of the wedge for a more rigid assembly of the stator and is also used to isolate the stator and rotor and form a closed space with the stator frame for the flowing of the cooling oil, which is not shown in Fig.\u00a02. The flux density has a direct impact on the performance and cost of the HTS generator, and that of air-gap, rotor yoke, and stator yoke after comparative analysis are chosen to be 1.4, 2.5 ~ 3.0, and 2.0 ~ 2.5\u00a0T, respectively. The heat load of the stator is chosen to be 6811 A2/(mm2.cm), the main parameters of the HTS generator are shown in Table\u00a01. Especially, the stator coils are divided into two columns and six rows corresponding to the twelve turns. The first row of the conductor is near the air gap, and the sixth row of the conductor is near the lamination, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002645_s00339-020-04238-2-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002645_s00339-020-04238-2-Figure11-1.png",
        "caption": "Fig. 11 Sample partition size and the hole spacing",
        "texts": [
            " It was observed that the surface residual tensile stress of a circular hole was relatively larger than that of a square hole. Considering the effect of square and circular holes on the surface residual stress, the analysis indicated that with the increase in aperture, the surface residual tensile stress first reached the maximum value and then gradually decreased. The thickness of base plate used in the experiment hb is 5\u00a0mm, and the size of test piece L \u00d7 B \u00d7 H is 50\u00a0 mm \u00d7 10\u00a0 mm \u00d7 4\u00a0 mm. Therefore, H > 2hb 3 ; The hole spacing of 3D printing sample is as displayed in Fig.\u00a011. So the sample is evenly divided into equal parts after consideration and analysis, and the size of each equal part is L1 \u00d7 B1 = 2.5 mm \u00d7 2 mm . It can be concluded that L1\u2215B1 = 1.25 . The minimum d\u2215B1 ratio of PLA sample printed is d\u2215B1 = 0.8\u22152 = 0.4 . With the increase in aperture, the stress concentration constant first increases to the highest value and then decreases when the d/B ratio is greater than 0.4 as presented in Fig.\u00a02. The surface residual tensile stress obtained in this experiment also increased with the increase in aperture, which first increased to the maximum and then decreased"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000683_tmag.2016.2601886-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000683_tmag.2016.2601886-Figure7-1.png",
        "caption": "Fig. 7. Open-circuit equipotential and flux density field distributions at aligned position.",
        "texts": [
            " Furthermore, for the PS-PMM-IIs, the PM thicknesses for 12S/10R, 12S/11R, 12S/13R and 12S/14R stator/rotor pole number combinations are 2.5, 2.6, 2.6, and 2.7mm (yellow marks) respectively since they not only have the maximum average torques but also have the approximately optimal demagnetization withstanding capabilities. B. Open-circuit Field Distribution The open-circuit equipotential and flux density field distributions for all machines at aligned position (negative daxis rotor position) are shown in Fig. 7. Obviously, for both PS-PMM-Is and PS-PMM-IIs, the flux loop of each coil belong to the same phase is completely independent within 12S/10R and 12S/14R while that is dependent within 12S/11R and 12S/13R, which is consistent with SS-PMM. Meanwhile, short flux path, which could results in lower MMF drop in the stator and thinner thickness of stator yoke, is also observed in all machines. Moreover, the saturation in PS-PMM-IIs are much heavier than those in SS-PMM and PS-PMM-Is, especially in the regions of stator yoke which close to the gaps between the adjacent PMs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002020_s12206-020-1030-6-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002020_s12206-020-1030-6-Figure3-1.png",
        "caption": "Fig. 3. Kinematic model of the finger mechanism.",
        "texts": [
            " Due to the symmetrical movement characteristic of the finger mechanism, it will be beneficial to achieve modular design and flexible assembly. In addition, the number of the components of the gripper could be greatly reduced and result in lower manufacturing cost. In the following sections, the kinematic modeling is introduced in Sec. 2. The effects of structural parameters on finger kinematic behavior is explored in Sec. 3. The effectiveness of the finger mechanism in bidirectional object grasping is validated experimentally in Sec. 4. And some conclusions and design suggestions are given in Sec. 5. As displayed in Fig. 3, a coordinate system is fixed at the midpoint of link AD. The length of link AB, BC, CD and AD are 1l , 2l and 4l , respectively. The length of the distal phalange 5l is described by the distance between point E and F. The point E is the midpoint of link BC. The rotation of link AB and BC is described by 1\u03b8 and 3\u03b8 , respectively. In the absolute coordinate system o-xy, the displacement of points B and C could be described as ( )4 1 1 1 1cos sin 2B ll l\u03b8 \u03b8\u239b \u239e= + +\u239c \u239f \u239d \u23a0 R i j (1) C B BC= +R R R (2) where ( ) ( )3 3 3 3cos sin "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001138_1.g003926-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001138_1.g003926-Figure3-1.png",
        "caption": "Fig. 3 Notation and conventions for CMGi.",
        "texts": [
            " \u00a7Note that eatt attitude error and ratt attitude command inFigs. 1 and2 only. D ow nl oa de d by C A R L E T O N U N IV E R SI T Y o n A ug us t 2 0, 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .G 00 39 26 II. Spacecraft Attitude Dynamics with Direct GimbalRate Request Thedynamics of a spacecraft SC consisting of a busB andn singlegimbal constant-speed CMGs in an arbitrary arrangement are considered. For i 1; : : : ; n, CMGi is composed of wheel Wi mounted on gimbal Gi as shown in Fig. 3. Vectors denoted by r \u21c0 and second-order tensors denoted by R ! are component free.\u00b6 Differentiation of component-free vectors and tensors is performed with respect to a frame; differentiation of resolved vectors and tensors does not require a frame. Let FI be an inertial frame, and let FB be a bus frame defined by the pyramidal CMG arrangement as discussed in Sec. V. The rotational attitude kinematics of the bus are described by Poisson\u2019s equation R !B\u2022 B\u2215I R ! B\u2215I\u03c9 \u21c0\u00d7 B\u2215I (1) whereB\u2022 denotes the frame derivativewith respect to FB, R ",
            " Assumption 3: For all i 1; : : : ; n, the center of mass ci ofWi is fixed in the bus. Assumption 4: For all i 1; : : : ; n, the direction of the axis i\u0302Wi is fixed in FGi . In particular, i\u0302Wi i\u0302Gi . Assumption 5: For all i 1; : : : ; n, the wheel Wi is inertially symmetric around i\u0302Wi . Assumption 6:For all i 1; : : : ; n, the frameFWi spins around i\u0302Wi at the constant rate \u03bdi > 0 relative to FGi . Assumption 7: For all i 1; : : : ; n, gimbal Gi is actuated by requesting its rate ui around j\u0302Gi relative to FB. Figure 3 and Assumptions 4 and 5 imply that the inertia matrix of Wi relative to ci resolved in both FGi and FWi is given by Ji \u225c J ! Wi\u2215ci jGi J ! Wi\u2215ci jWi 2 4 \u03b1i 0 0 0 \u03b2i 0 0 0 \u03b2i 3 5 (6) where \u03b1i is the moment of inertia of Wi around the spin axis i\u0302Wi i\u0302Gi , and \u03b2i is the moment of inertia around the remaining axes of FWi and FGi . Therefore, J ! Wi\u2215ci jB Oi J ! Wi\u2215ci jGi OT i OiJiOT i (7) and the 3 \u00d7 3 orientation matrix Oi is defined by Oi \u225c OB\u2215Gi R ! Gi\u2215BjGi R ! Gi\u2215BjB (8) where R ! Gi\u2215B is the rotation tensor that rotates FB into FGi ",
            " Wi\u2215ci u \u21c0\u00d7 i \u03c9\u21c0B\u2215I u \u21c0\u00d7 i J ! Wi\u2215ci \u03c9 \u21c0 Wi\u2215Gi u \u21c0 i J ! Wi\u2215ci u \u21c0 B\u2022 i J ! Wi\u2215c\u03c9 \u21c0 B\u2022 B\u2215I (A23) B. Resolving the Equations of Motion in FB Resolving Poisson\u2019s equation [Eq. (1)] in FB yields _RB\u2215I d dt R!B\u2215IjB R !B\u2022 B\u2215IjB R ! B\u2215IjB\u03c9 \u21c0 B\u2215Ij\u00d7B RB\u2215I\u03c9 \u00d7 B (B1) where RB\u2215I \u225c R ! B\u2215IjB R ! B\u2215IjI; \u03c9B \u225c \u03c9 \u21c0 B\u2215IjB (B2) Resolving Eq. (A1) in FB yields H \u21c0 B\u2215c\u2215IjB JB\u03c9B (B3) where JB \u225c J ! B\u2215cjB (B4) Furthermore, resolving Eq. (A18) in FB yields H \u21c0 B\u2022 B\u2215c\u2215IjB d dt H\u21c0B\u2215c\u2215IjB JB _\u03c9B (B5) where _\u03c9B \u03c9 \u21c0 B\u2022 B\u2215IjB (B6) Using Fig. 3 and Assumptions 4, 5, and 6 to resolve the angular velocity of FWi relative to FGi in FGi yields \u03c9 \u21c0 Wi\u2215Gi j Gi \u03bdie1 (B7) where \u03bdi > 0 is the angular rate of Wi around i\u0302Wi relative to FGi . Similarly, the gimbal angular velocity u \u21c0 i defined by Eq. (A8) and the gimbal angular acceleration u \u21c0 i Gi\u2022 of FGi relative to FB resolved in FGi are given by D ow nl oa de d by C A R L E T O N U N IV E R SI T Y o n A ug us t 2 0, 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .G 00 39 26 u \u21c0 ijGi uie2; u \u21c0 i Gi\u22c5 jGi d dt u\u21c0ijGi _uie2 (B8) where ui is the gimbal-rate request for CMGi, and _ui is its derivative. Because ui is the spin rate of gi, the corresponding gimbal angle \u03b8i is defined such that _\u03b8i ui (B9) It thus follows from Eqs. (A8) and (B8) that \u03c9 \u21c0 Gi\u2215BjGi _\u03b8ie2 (B10) Defining the gimbal angle vector \u03b8 \u0394 \u03b81 \u00b7 \u00b7 \u00b7 \u03b8n T yields _\u03b8 u, where u \u0394 u1 \u00b7 \u00b7 \u00b7 un T \u2208 Rn is the control input. Furthermore, Fig. 3 and Assumptions 4 and 5 imply that the inertia matrix ofWi relative to ci resolved in both FGi and FWi is given by Ji \u225c J ! Wi\u2215ci jGi J ! Wi\u2215ci jWi \" \u03b1i 0 0 0 \u03b2i 0 0 0 \u03b2i # (B11) where\u03b1i is themoment of inertia ofWi around the spin axis i\u0302Wi i\u0302Gi and \u03b2i is the moment of inertia around the remaining axes of FWi and FGi . Next, note that u \u21c0 i B\u2022 u \u21c0 i Gi\u2022 \u03c9 \u21c0 Gi\u2215B \u00d7 u \u21c0 i u \u21c0 i Gi\u2022 (B12) Resolving Eqs. (B7), (B8), and (6) in FB and using Eq. (B12) yield \u03c9 \u21c0 Wi\u2215Gi j B \u03bdiOie1; u \u21c0 ijB \u03c9 \u21c0 Gi\u2215BjB uiOie2; u \u21c0 i B\u2022 jB u \u21c0 i Gi\u2022 jB _uiOie2 (B13) where J "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002077_j.mechmachtheory.2020.104209-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002077_j.mechmachtheory.2020.104209-Figure1-1.png",
        "caption": "Fig. 1. Structure of an Exechon-like PKM: (a) PKM, (b) limb assemblage.",
        "texts": [
            " In Section 5 , a parameter sensitivity analysis is carried out to identify the most influential geometric variables and to optimize the geometrical tolerance allocations of the over-constrained PKM by considering parameter uncertainties. Finally, some conclusions and remarks are presented in Section 6 . From a topological point of view, an Exechon-like PKM claims the characteristics of tripod-arrangement, over-constrained structure, three prismatic actuated limbs as well as six passive joints with thirteen equivalent single degree of freedom (DOF) joints [39] . As depicted in Fig. 1 (a), an Exechon-like PKM mainly consists of a moving platform, a fixed base and three prismatic actuated limb assemblages, of which three limbs are attached to the moving platform and the fixed base by two passive joints (located at A i and B i , i = 1, 2, 3). Each individual prismatic actuated limb assemblage includes two passive joints, a limb body, a lead screw assemblage and a servo motor as shown in Fig. 1 (b). As can be seen from Fig. 1 , when the passive joint at A i ( i = 1, 2, 3) performs as a revolute joint (R joint) and the passive joint at B i ( i = 1, 3) performs as a universal joint (U joint) while the passive joint at B 2 acting as a spherical joint (S joint), the Exechon-like PKM will claim a topology of 2U P R-1S P R ( P : prismatic actuated joint). In other words, it constructs an Exechon PKM. By taking different joint forms at A and B ( i = 1, 2, 3), one may construct different types of PKM with similar i i topological structure with the Exechon",
            " With the above Cartesian coordinate settings, the transformation matrixes between the body-fixed frames P-uvw, A i -x i y i z i and the global reference frame O-xyz can be derived as O P : T rans ( P \u2212 u v w \u2192 O \u2212 xyz ) (1) O Ai : Trans ( A i \u2212 x i y i z i \u2192 O \u2212 xyz ) (2) where O P and O Ai denote the transformation matrixes of P-uvw and A i -x i y i z i with respect to O-xyz , respectively, which can be referred to our previous publication [40] . The Exechon-like PKM is simplified into an equivalent kinetostatic model, in which all passive joints are treated as virtual spring units with lumped-parameter method while the limb body is modeled as an elastic spatial beam through the finite element method [41] . The Exechon-like PKM shown in Fig. 1 (a) can be simplified into the equivalent kinetostatic model as depicted in Fig. 2 . As depicted in Fig. 2 , the kinetostatic model of the Exechon-like PKM is assumed with the modular components that mainly consist of the elastic joints and limbs as well as the rigid platform and base. The passive joints are represented by a set of virtual spring units with equivalent translational/torsional stiffness constants of k L Ai / k A Ai and k L Bi / k A Bi , whose corre- sponding linear/angular elastic deformations are denoted as x L Ai / x A Ai and x L Bi / x A Bi , respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002342_1475921720916227-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002342_1475921720916227-Figure2-1.png",
        "caption": "Figure 2. Test rig for validation of proposed method: (a) test rig; (b) roller speed detection without outer raceway; (c) roller speed detection with outer raceway; (d) calibration of roller speed detection with Hall sensor.",
        "texts": [
            " The identified IMFs (signal IMFs) contain the rotational frequency of not only the roller but also the inner raceway ~f\u00bc fr; 2fr; . . . , nfr; fm\u00f0 \u00de \u00f011\u00de Finally, the cage slip can be obtained by extracting the feature frequency (fr and fm) of reconstructed signal. Table 1 is the identified framework of cage slip. The detection of roller speed is a key mask for the study of the cage slip. The speed mainly depends on the detection of the WMF of the roller. Thus, in this section, a test rig, as shown in Figure 2, is designed and developed to verify whether or not the proposed method can detect magnetic field information of roller. The test rig consists of motor, governor, rotational axis, and supporting mechanism. Here, the rotational axis is a plastic pipe instead of conventional rigid products. The aim is that the plastic pipe is consisted of no permeability material. The material cannot generate the WMF. It can eliminate the influence of rotational shaft. The following is the verification process: 1. A roller is installed on the rotational shaft. The WMF sensor is used to detect the magnetic field of the roller as shown in Figure 2(b). 2. The outer raceway shields the roller as shown in Figure 2(c). The WMF sensor is also used to collect the magnetic field information. The feasibility of proposed method is verified by comparing with magnetic field information in both cases. 3. The circular magnet (Nd-Fe-B magnet) is installed on the rotation axis as shown in Figure 2(d). The Hall sensor is used to detect the magnetic field of circular magnet to calibrate proposed method. The detected signal is plotted in Figure 3. The roller signal, which is not shielded by outer raceway, is plotted in Figure 3(a). The roller signal, which is shielded by outer raceway, is plotted in Figure 3(b). Figure 3(c) plots the detected signal of circular magnet with Hall sensor. The comparison of detection signal with WMF sensor is also plotted in Figure 4. Findings show that the detected signal with WMF sensor is phase reverse in the both cases as shown in Figure 4(a)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003566_pime_proc_1945_153_010_02-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003566_pime_proc_1945_153_010_02-Figure13-1.png",
        "caption": "Fig. 13. Bogie Suspended by Two Parallel Chains",
        "texts": [],
        "surrounding_texts": [
            "If the maximum angle of the chain is denoted $y 8, then the maximum displacement of point H = I sin 8; and the acceleration f at the end of the swing (rw2) is given by f = lsin e($)' Let w be the weight of bogie and d the distance between the supporting chains. Then the restoring torque at the end of the swing is = (+w sin0)xd If the acceleration of hook H is f, the angular acceleration of the bogie is 2f radians per sec. per sec. d Torque = moment of inertia x angular acceleration W 8n2 p c x d x sin 0 =-kzxIsin e x = g gd2tZ 165~~1 4n I k2 = -so And k = - dt J\" k = 15.3 x 2.25JZ7 _ - 3.40 feet For motor bogie, 4n 13.92 X 2.76 __ For trailer bogie, k = 4n 4% = 3.79 feet For driving trailer bogie, k = _ - 4n Coimmunicatioiis Captain E. F. BLAKEMORE, B.Sc., R.E.M.E., G.I.Mech.E., wrote that since the paper must be of profound interest to all students of engineering, it was a great pity that the author had omitted the history of wheel coning; for instance, he did not mention that parallel treads had bee3 tried by several railways without success; nor did he mention that railways had tried, and ceased to use, wheels coned 1 in 15, 1 in 24, and 1 in 32, or that one railway many years ago tried various conings and parallel treads on the same train at the same time. Perhaps he would give a bibliography. No information was given of the physical properties of the wheels and rails, or the condition of the track-all inseparable from tyre performance and wear, though the maximum amplitude of the wavs motion given in the tables implied that the rails were well worn. It was thought by some engineers that the rail contour changed but little during life, while others held that it did not matter what the original contour was, since it did not last long. He would like to know the author's opinion on that matter. When rails were new, the track could be represented by two circles 24 inches diameter upon which the various forms of wheel ran. With parallel treads perfectly matched, the wheels could be represented as a cylinder rolling along the tops of the circles obviously without horizontal restraint, since the points of contact were exactly on the vertical diameters (Fig. 14). That lack of restraint explained why wear on parallel treads was not a pure radius as on the majority of worn coned treads. If the wheels were imperfectly matched they would arrange them- selves on the circles differently; and when running one wheel of each pair would drag-but which one? Did it not depend whether it was a non-driver or a driver? For whereas a non- Fig. 14. Path of Wheel Treads Perfectly and Imperfectly Matched driver would have the same tractive effort applied to its tread, no matter what the diameter, the driver had a tractive effort inversely proportional to the diameter; so it would seem that on non-drivers the smaller wheel dragged, while on the drivers at CAMBRIDGE UNIV LIBRARY on June 4, 2016pme.sagepub.comDownloaded from 36 COMMUNICATIONS ON R I D I N G AND WEARING QUALITIES OF RAILWAY CARRIAGE TYRES the larger wheel would tend to drag. Had the authormy definite proof that small wheels always dragged? If the wheels were coned equally and the diameters were the same they would ride centrally on the circles, unless disturbed by a side blow, when lateral creep would take place on both wheels alternately. Should the wheels be of the same coning but of different diameters, they would tend to centralize with an inclined axle (see Fig. 15) with consequent flange wear. It seemed obvious that the wear of treads and flanges depended not only upon the original shape but also upon the restrictions imposed by the bogie. Here again there was a difference of opinion between various authorities; but was it not true that in the majortiy of cases wheels coned 1 in 20 wore to a 12-inch radius on the tread? He was struck by the five-figure dimensions of the wheels, which made him wonder whether the last figure was accurate, especially when taken on a coned wheel. And the closeness of the pairing suggested a care which, with the accompanying high cost of turning, was undesirable, if not unattainable, in the ordinary shop. Perhaps the author would explain how he measured his diameters. Assuming the figures true, the results bore out the assumption that, for cylindrical treads at any rate, the larger driving wheels wore more than the small ones, while the smaller non-driving wheels wore more than the larger ones. That was not so marked with the 1 in 100 coning, and was probably due to the self-adjusting properties of the wheel. The trailer wheels did follow the usual rule, that is, the smaller wheels wore more. If the author did make measurements other than those shown in his tables, perhaps he would state whether the trailer bogies \u201ccrabbed\u201d, thus causing wheels at the extreme and opposite corners to wear more on the flahges than the intermediates. He suggested that except for \u201cseparation\u201d of flange and tyre wear the wheel and flange wear should not be measured by the comparison of new profile with worn profile, but of initial profile with profile after re-turning, because the amount of material removed by wear .was approximately 25 per cent, while that removed by turning was about 75 per cent. For that reason he did not agree with the author\u2019s figures given in Table 16, p. 33 of the paper, nor with the statement that profiles could be brought to a common basis for discussion. Generally, what really mattered was (1) good riding of the. coach, ;ind (2) maximum mileage out of tyres from replacement to replacement. The latter condition meant a balance of wear between flange and tread since, when re-turning, a wear of 0.1 inch on the flange meant much more than a wear of 0.1 inch on the tyre. Finally, would, the author explain why the gauge measured the flange wear at a distance of inch from the top of the flange? Whilst he thought that the method adopted by the author to obtain the values of lateral oscillations was very ingenious, and whilst in general he agreed with his conclusions, he felt that it was rather presumptuous to suggest that a 1 in 100 coning would cure, or partially cure, all the ills to which rolling stock was prone when on the strength of results obtained with only 104,000 miles\u2019 running on electric stock, as compared with the millions of miles run by all kinds of stock. He hoped that the paper was only an interim report. Mr. W. A. J. DAY, A.M.I.Mech.E., wrote that tyre life on the South African Railways was formerly determined, in the majority of cases, by flange wear. In later years, due to gradual reduction or elimination of severe curvature, together with improvement in design of vehicles, flange wear had been reduced appreciably. By these means, tyre life had been extended and had been accompanied by an increasing proportion of cases in which tread wear was the determining factor. However, flange wear was still the greater factor. The minimum permissible thickness of flange adopted by the South African Railways for tyres was 3 inch at half depth (& inch). The main danger of sharp flanges on the South African Railways was the tendency to open or foul facing points. The reduction of strength of tyre at root of flange due to wear was not. considered to be an important factor. Tyres were inclined 1 in 20, and rails were similarly canted. About the year 1936, cylindrical tyres were tried on.some electric motor coaches running between Capetown and Simonstown. Ascertainable results were indeterminate, and the experiment was not continued. Those motor coaches were fitted originally with a centring device controlling the lateral movement of the bogie bolsters. Owing to rather violent oscillations occurring on the coaches, the controlling device was removed and improved riding qualities resulted. It was considered that the comparatively low centre of gravity of motor coaches aggravated any tendencies to oscillation and flange wear. Steeply coned tyres were considered to increase lateral oscillations, particularly at high speeds. Conditions in South Africa were not comparable with those in Britain. Curvature of track was more severe and more prevalent. Speeds were lower (55 m.p.h. was the official maximum), but tended to increase. Under existing South African conditions, cylindrical tyres would probably require more frequent attention and re-turning without corresponding advantages, although some reduction in steepness of the coning appeared to be worth consideration. He also felt that there were several other factors having important bearing on tyre wear and riding qualities of vehicles. For instance, the maintenance of reasonably correct parallelism of axles was important. The use of narrow horns and axlebox guide surfaces liable to undue wear, as well as types of axlebox bearings and bearing seats having wide fitting tolerances (thus permitting appreciable divergence from parallelism), was probably responsible for avoidable tyre wear and unsatisfactory riding of vehicles. If some diversion from the subject-matter of the paper was permissible the following were some of the principal methods adopted by the South African Railways during the past twenty years to reduce tyre flange wear and to improve the riding qualities of vehicles :- (1) Use of high-tensile (63-69 tons per sq. in.) tyre steel accompanied by adoption of the Gibson ring tyre fastening in place of stud fastening. Tyre failures had been reduced almost to a negligible quantity. (2) Adoption of vertical swing links for coach and motor coach bogie bolsters, in place of inclined links. (3) Use of bogie centres having flat wearing surfaces, oilimmersed where practicable, in place of spherical or curved centres. (4) Locating the bogie side. bearers nearer to the bogie centre and arranging for their wearing surfaces to be oilimmersed. Mr. C. E. FAIRBURN, M.A., M.I.Mech.E., wrote that it might be of interest to record the origin of the experiments on the London, Midland and Scottish Railway which were the subject of Mr. Newberry\u2019s paper. In 1935 Sir Harold Hartley and the writer, when travelling in the United States, visited the Chicago at CAMBRIDGE UNIV LIBRARY on June 4, 2016pme.sagepub.comDownloaded from COMMUNICATIONS ON R I D I N G AND WEARING QUALITIES O F RAILWAY CARRIAGE TYRES 37 Rapid Transit Company and were very kindly shown the film records and given a full account of the tests with cylindrical tyres which had been made on vehicles using that company\u2019s lines. In the London, Midland and Scottish Railway tests the question was considered largely in its relation to the riding qualities of the stock, but the Chicago tests arose from considerations of track maintenance. During the spring thaw, after the severe winter frosts experienced in that region, very heavy expenditure was incurred in track maintenance and re-alignment. Experience showed that that expenditure was substantially reduced when cylindrical treads were substituted for the 1 in 20 taper profile previously in use. As the lines ran through the city, there were many sharp curves, and flange wear was rapid; but it was found that there was no appreciable difference between the two profiles in that respect. In any case, the saving in civil engineering expenditure was much more than sufficient to compensate for any possible increase in the cost of tyre re-turning. The question was discussed also with the Illinois Central R.R. and the Chicago, Burlington and Quincy R.R. These companies were interested primarily in the riding of the stock; the former had been using cylindrical treads on multiple-unit electric vehicles with success for two years or more, and the latter had-adopted them for some axles on the Burlington Zephyr train, but neither company could make a definite pronouncement about their wider adoption on ordinary steam-hauled stock. With this information about American experience available, it was decided that tests should be made in this country, and the electric stock on the Liverpool-Southport line was very suitable for the purpose. Quite apart from the records given in the paper, there was no doubt at all from personal observation that the cylindrical tyres gave markedly more comfortable riding than the standard 1 in 20 taper profile. There was not a great difference between the cylindrical tyre and the 1 in 100 taper; but what advantage there was, lay with the former. That was to be expected from theoretical considerations. The wavelength L of the movement of a single axle was given by L = 27~); where r denoted the wheel radius, 2b the gauge, and T the tangent of the coning angle of the tyres.* If T = 0, i.e. if the tyres were parallel, the oscillatory movement ceased. The paper made it clear, however, that the cylindrical tyre presented a difficulty in that the diameters of both wheels on an axle must be matched to closer limits than was practicable with normal machining methods, whereas with the 1 in 100 taper, there is less difficulty with the limits which could be allowed, owing to the ability of the taper tread to compensate for some difference in diameter. The damage to the track and the rough riding investigated in America were probably due largely to bogie \u201chunting\u201d. There could be little doubt that that elusive effect was the result of forced oscillations set up by the periodic movements of the wheels, and it was not experienced during the London, Midland and Scottish Railway tests. It was, however, interesting to observe that the period of the wheel movements was about the same as the usual frequency of hunting, i.e. about three per second. A point calling for greater emphasis was the interaction between the tyre and worn rails; it was general knowledge That when rails and tyres became worn to approximately the same profile, bad riding resulted. An improved performance with a new profile, or for that matter a re-turned tyre of the existing profile, might, therefcre,\u2019 be due-at least partially-to the fact that the tyres and the rails had different shapes, and not to the merits of the particular tyre profile as such. That made it difficult to decide beforehand what result would be obtained if the whole of the company\u2019s stock were changed to a taper of 1 in 100. In time the rails would become worn to the new profile, and it might be that the riding would not then be appreciably better than that obtained at present. The suggestion that a railway shodd operate with two different profiles, so that the rails * DAVIES, R. D., 1938-9, JI. Inst. C.E., vol. 11, p. 231, \u201cOscillation of Railway Vehicles\u201d. never \u201csettled down\u201d to either, seemed, therefore, to be worth serious consideration, and might well be explored further. The paper introduced a \u201criding factor\u201d in evaluating the results. That factor had dimensions of time divided by length; and while it had the merit of giving a figure which approximated to what was observed in practice, its true significance seemed to be doubtful, and it should not be taken as more than an indication of performance. Mr. W. S . GRAFF-BAKER, M.I.Mech.E., wrote that the author referred to similar work carried out on the riding qualities of cylindrical tread wheels by the Chicago, North Shore and Milwaukee Railroad. There seemed to be one fundamental difference between the experiments on this railroad and those on the London, Midland and Scottish Railway in that the American experiments were conducted on a horizontal-top rail whereas the London, Midland and Scottish experiments were carried out on an inclined rail, It would be interesting to know what bearing the difference would have on experimental results. On p. 33 of the paper, at the end of section (1) of Part I11 the author referred to some improvement in riding factor that was sometimes obtained by the tread of the wheel wearing hollow. Did not that indicate a possible line of research in the direction of making at least one tyre of the two on an axle hollow in the tread in the first instance? Under the heading of \u201cTread Wear\u201d in section ( 2 ) of the test results (pp. 33-34) reference was made to the tread wear on a driving wheel being greater than that on a non-driving wheel. Surely that was related to two factors : (1) the tractive effort exerted at the tread which might accentuate differential slipping, and (2) the considerably greater weight carried by the driving wheel. The material of the brake blocks had a considerable bearing on tyre wear, whether on the tread or on the flange. The London Passenger Transport Board used cast iron blocks in some cases, with a very considerable flange wear, flange wear in fact being the governing factor in considering the need for re-turning the tyres. Some blocks engaged the flanges, others did not-the results did not appreciably differ. On the District Line, using cast iron blocks, with trains stopping about every + mile, the mileage between tyre turnings on motor wheels was 34,400 and on trailer wheels 60,200. On the Northern Line, using nonmetallic blocks, the tyre turning was governed by tread wear alone, and the mileage between turnings was 142,400 for motor wheels and 167,700 for trailer wheels. The Northern Line trains stopped a t similar intervals to the trains on the District Line. The effect of a non-metallic block was to impart a fairly well polished surface to the tread and therefrom to the rail, reducing rail wear very substantially, particularly on curves, due, it was thought, to permitting differential slipping action to take place without undue friction. Mr. F. C . JOHANSEN, M.Sc. (Eng.), M.I.Mech.E., wrote that apart only from establishing, by full-scale experiment, the major fact that reduced coning improved riding, the most farreaching outcome of the author\u2019s investigation had been to emphasize the useful part played by coning in counteracting effects of inevitable small differences between the diameters of mated wheels. The preferential wear of cylindrical profiles, due to such differences in diameter, seriously offset their good qualities in respect of riding; and it was that factor that prompted the suggestion-originally made by Mr. T. M. Herbert, M.A., M.I.Mech.E., Research Manager of the London, Midland and Scottish Railway-to compromise with 1 in 100 coning, which had proved satisfactory from the dual standpoints of good riding and resistance to wear. For low-speed passenger stock and freight wagons, where tyre profiles conducive to good riding were less important than for express passenger services, it might be that some coning intermediate between 1 in 100 and the present standard 1 in 20, would be found by experience to yield the best all-round results. But while some general change of tyre profile might not be out of the question, any corresponding departure from the present 1 in 20 inclination of rails seemed too remote for contemplation. It therefore became of interest to consider one or two practical questions that arose from the assumption that wheels coned 1 in 100 were to run on rails canted at 1 in 20. at CAMBRIDGE UNIV LIBRARY on June 4, 2016pme.sagepub.comDownloaded from 38 COMMUNICATIONS ON R I D I N G AND WEARING QUALITIES O F RAILWAY CARRIAGE TYRES The 1 in 100 coned tyre profiles .used in the tests were a simple development of the present standard 1 in 20 coned profile, derived by joining a tread, tapered 1 in 100, tangentially to the root of the standard flange. Some consequences of running that modification on a rail canted at 1 in 20 to the vertical were shown, for new rail and tyre profiles, by Fig. 16 on which the standard 1 in 20 coned profile also appeared for comparison. The 1 in 100 coned profile, riding on the outer part of the rail, was lifted higher than the 1 in 20 coned profile, with the results, not only that the rail made contact farther down the flange (as mentioned in the paper) but also that the tip of the flange projected a somewhat shorter distance below rail level, and that the clearance between the rail and the flange when the wheels were central was somewhat increased. The present standard flange (contour A, British Standard Specification No. 276,1927) presumably had the optimum depth and shape respectively to guard against derailment and allow the correct lateral play of a pair of wheels within the rail gauge. The flange associated with 1 in 100 coned tread should, therefore, extend to the same radial distance below rail level, and should allow the same minimum transverse play. Those features could be achieved, as indicated in Fig. 16, by a flange profile of the shape ABCD whereby, in effect, the present standard 1 in 20 coned flange in correct relation to the rail was combined with the 1 in 100 coned tread, the only new portion of profile being the short length AB in the throat of the flange which, incidentally, did not make contact with the rail so long as the profiles were unworn. That suggested flange for 1 in 100 coned tyres extended about inch farther below rail level than the simple modified profile used in the tests. The common tangent where it made contact with the rounded corner of the rail was inclined at about 50 deg. to the horizontal, as compared with about 55 deg. for the profile used in the tests, and 37 deg. for the 1 in 20 standard flange profile. Those angles would, of course, be markedly altered by a small amount of wear. inch, the minimum lateral clearance between rail and one tyre, the suggested profile increased equally the maximum clearance between the other tyre and its associated check rail. An unworn cylindrical or 1 in 100 coned tread made contact with the rail head at about 3 inch to the outer side of the \u201cvertical\u201d axis of the rail. That was confirmed experimentally by the appropriate cinematograph pictures which showed a gap between the rail and the tread towards the flange side, with turning marks on the inner part of the tread still visible after considerable mileage. The consequent eccentricity of loading must accentuate stresses in certain parts of the rail, and also introduce an overturning moment, acting in the same sense as lateral flange forces, which must be resisted by the keys. Neither of those effects was likely to be serious, but would no doubt be worth investigation, or at least consideration, if 1 in 100 coning was introduced to any wide extent. The effect of the upward rail By reducing, to the extent of about reaction on the wheel being horizontally slightly nearer the downward load on the journal (as compared with the case of the standard 1 in 20 coned profile) was to reduce the bending moment in the axle and was therefore beneficial. Mr. H. N. LINTON, B.Eng., A.M.I.Mech.E., wrote that the test particulars and data, also the method of transcription from film to graphical record, were extremely interesting and emphasized the need for improvement by research instead of by guess-work or \u201ccommonsense\u201d methods. The summaries in Tables 10-13, pp. 31, 32 of the paper, in spite of many inconsistencies, generally supported the author\u2019s conclusions. It should be noted, however, that the tests described took place at relatively high speeds (55-65 m.p.h.) on track which was mainly straight, level, and free from junctions with their attendant check rails. Perhaps the author would say if any tests had been conducted at a range of speeds over \u201cmixed\u201d track and, if so, whether the results supported his present conclusions. The results of the tests on the standard A.R.L.E. (1 in 20) profile tyres were rather disturbing at first glance, the period of oscillation dropping to less than one-half of the original value after 20,000 miles, whilst the \u201criding factor\u201d dropped to something approaching one-half of the original value. He himself doubted the implication that a ride in a vehicle with worn tyres was twice as uncomfortable as that in a vehicle with new tyres, particularly in view of the fact that, on the author\u2019s figures, the amplitude of oscillation was greatly reduced and, therefore, the tendency for the flanges to crash against, and even ride, the rails was likewise reduced. There was one other point which the author might clarify. The wheel diameters varied throughout the tests as a result of turning the tyres or fitting new tyres and, consequently, the heights of the centres of gravity of the vehicles varied also. Did the author consider that the latter variation, in the region of $ inch, had any bearing on the results? Mr. C. F. DENDY MARSHALL, M.A., wrote that one could not be surprised that wheels with coned tyres followed a sinuous path. A pair of wheels on an axle, of different diameters, would tend to travel in a circle, the centre of which was on the line of the axle, beyond the smaller of the two. With coned tyres, if for any reason, at a given moment the system was displaced to the left with reference to the rails, the wheel on that side was running on a larger diameter and the system would bear to the right. Its momentum would carry it beyond the centre, and there would be a swing back to the left. If that was the whole explanation, the waves would be gradually damped down, until they were started again by some fresh impulse received from the rails, or perhaps through the coupling. Moreover, the phenomenon should not occur, or at all events it should not occur regularly, in the case of cylindrical tyres. There must therefore be some other cause, which at present seemed to be hidden. The wear of tyres on carrying wheels seemed to be due to three main causes : (1) side-slip due to lateral oscillation; ( 2 ) longitudinal slip caused by difference of diameter (always present with coned wheels except when exactly in the correct lateral position); and (3) braking. The first could be reduced by cutting down the side play allowed, which was a legacy from the days of carriages with rigid wheelbases. The second could only be eliminated by arranging for one or both wheels to be free to turn on the axle. I t was impossible to estimate the amount of longitudinal slip that occurred with coned wheels, but it was easy to calculate it in the case of cylindrical ones. It was far greater than was generally realized. Imagine a pair of cylindrical wheels running from London to Edinburgh. If they weremominally 3 ft. 6 in. in diameter, the circumference was 11 feet, and they would make 192,000 revolutions. The author gave 0.01 inch as a reasonable limit of the difference of diameter, which was 0.0314 inch on the circumference. On the journey, therefore, there would be a slip of 6,040 feet, or well over a mile. The third cause of wear could be removed by providing special drums or disks for braking. With regard to the wear of driving wheels, it could be taken that the better the adhesion, the less the wear. For some time it had been realized that the adhesion diminished with speed. at CAMBRIDGE UNIV LIBRARY on June 4, 2016pme.sagepub.comDownloaded from"
        ]
    },
    {
        "image_filename": "designv11_14_0000683_tmag.2016.2601886-Figure24-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000683_tmag.2016.2601886-Figure24-1.png",
        "caption": "Fig. 24. Rotor lamination with bridge for 12S/10R PS-PMM-I.",
        "texts": [
            "75 TRR (mm) 5 RISI (mm) 10.4 TPM (mm) 4 \u03b8OSTB (\u00b0) 8.12 \u03b8PM (\u00b0) 30 \u03b8OSTT (\u00b0) 4.94 TBRI (mm) 0.5 TOSTTB (mm) 3.0 stator (12 PMs), outer stator (12-pole) and modular rotor (10- pole). Hence, the parameters of prototype machine as shown in Table VI are different from the previous globally optimized parameters as shown in Table I. Moreover, for easing the fabrication, 10-pole modular rotor is mechanically connected by the lamination bridges (TBRI = 0.5mm) in the inner side, as show in Fig. 23 (d) and Fig. 24. Fig. 25 shows the predicted and measured phase backEMFs at rated speed (400rpm). The measured fundamental value is ~4% less than the prediction, which is mainly due to the end effect in 25mm stack length machine. The predicted and measured open-circuit cogging torque waveforms are shown in Fig. 26. It can be seen that the measured peak to peak value is slightly larger than the FE prediction. This difference is acceptable when considering the measurement error and assembling tolerance. Fig. 27 shows the waveforms of static torque against with rotor position at five different 0018-9464 (c) 2016 IEEE"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001771_s12239-020-0088-6-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001771_s12239-020-0088-6-Figure1-1.png",
        "caption": "Figure 1. Vehicle layout.",
        "texts": [
            " Considering that the driving/braking force affects the roll motion of the vehicle, how to integrate the yaw stability and roll stability of the independent drive commercial vehicle is an urgent problem to be solved. At present, there are relatively few studies in this aspect. The motor has a fast response and precise control, but its adjustment capability is limited compared to the braking torque of a conventional brake system. How to make full use of the advantages of motor and hydraulic/ *Corresponding author. e-mail: bitev@bit.edu.cn pneumatic braking to improve vehicle stability is a direction of independent drive dynamics control research. As shown in Figure 1, the electric bus studied in this paper is equipped with an in-wheel motor drive system on the rear axle, and the two motors can be controlled separately. As the height of the center of gravity of the bus is increased and the wheelbase is narrowed, the anti-roll ability of the vehicle is reduced. Therefore, an integrated control strategy for vehicle yaw stability and roll stability is proposed. On the basis of the two-layer hierarchical structure mentioned above, the stability monitoring layer is added to the control strategy to make real-time decision on the vehicle stability control mode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002430_tie.2020.3014570-Figure19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002430_tie.2020.3014570-Figure19-1.png",
        "caption": "Fig. 19 3-D structure of the slotted LATM for illustrating the assembly process of the wounded tooth.",
        "texts": [
            " In addition, when the winding current is 10 A, the peak-to-peak torque ripple is 0.071 for the LATM with curved teeth, which is less than the corresponding 0.0732 for the straight teeth one. Therefore, the torque profile of the curved one has more symmetric distribution within the operating range [-15\u00b0, 15\u00b0], and also has a smaller torque ripple. IV. PROTOTYPE AND EXPERIMENTAL RESULTS A prototype with the curved stator teeth is manufactured shown in Fig. 18. To illustrate the assembly process of wounded tooth, a 3-D construction of the LATM is shown in Fig. 19. In order to improve the coil space factor and make the process of installing windings easier, the stator core is divided into three segments: a non-wounded stator core portion shown in Fig. 18(c) and two wounded teeth portions shown in Fig. 18(a). The coils can be wounded on the stator tooth firstly, and then the wounded tooth can be assembled into the non-wounded stator core. The test bench is established as shown in Fig. 18(d). Fig. 20 shows the comparison of output torque by the experimental and FEA results for the slotted LATM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003715_s0301-679x(98)00106-6-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003715_s0301-679x(98)00106-6-Figure2-1.png",
        "caption": "Fig. 2. Optical profile projector scheme.",
        "texts": [
            " Moreover, a comparison between contact pressures computed both numerically and analytically, using photoelastic data, has been carried out. The seal under test is a commercial lip seal used for sealing on pneumatic cylinder rods. The geometry and main dimensions are shown in Fig. 1. The seal has large dimensions in order to have a better visualisation of isoclinic lines and isochromatic fringes. The cross section has been measured by means of an optical profile projector (Optics Jena MP320); to this aim a thin segment of the seal has been cut and measured on the projector display with a 20\u00d7 magnification image (Fig. 2). The seal material is nitrile rubber NBR with hardness of about 71 Shore A. The material mechanical stress\u2013 strain characteristic has been determined performing an uniaxial tensile and compression test. The seal mechanical characteristics are needed for the finite element model using the experimental data directly as input. The tensile test has been performed using a dumbell specimen, according to ASTM D412-83 standards; in the compression test a cylindrical specimen (undeformed height Lo=10 mm, diameter d=10 mm) compressed between two well lubricated plates has been used"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure22.2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure22.2-1.png",
        "caption": "Fig. 22.2 Exoskeleton load lifting algorithm",
        "texts": [
            " It is necessary to observe a number of rules for effective load lifting: keeping of the operator\u2019s back in the correct position (so that the spinal column physiological curves remaining), maximizing of the legs usage during lifting process, and keeping person\u2019s arms outstretched [17]. During the load lifting process, three groups of joints located in the sagittal plane are mainly involved: the hip, knee, and ankle joints [18, 19]. The hands also perceive the load, but to reduce it, flexible connections redistributing the load on the exoskeleton power frame can be used. The load lifting process scheme with the usage of exoskeleton is shown in Fig. 22.2. In this case, the exoskeleton can assist a person when lifting a load, in particular, compensating for a part of the hip and knee joints torques. Note that in this case, the exoskeleton can be considered as a three-link mechanism in which the force load is applied at the sling attachment point O4 (Fig. 22.3). The load force FL = MLg acts on the three-link system in such a way: at the point O4(ML\u2014mass of the cargo). The links are affected by given moments, which can be created by either passive (springs, brakes, etc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001457_iros40897.2019.8968128-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001457_iros40897.2019.8968128-Figure5-1.png",
        "caption": "Fig. 5: Multi-leg coordination patterns. Schematics to represent and visualize the different patterns. In (a) Cruse, in (b) CPG and the leg numbering, and in (c) CPG local.",
        "texts": [
            " The patterns are also implemented as behaviours that are complementary, so that only one pattern is active at a time. A pattern class takes care of the motivation, activity and reflection for each pattern. The pattern class also detects if the robot is in a state of walking or not. The robot is in walking state if a direction is given and a pattern is active. The motor commands sent to the legs are only published to the actual joints only if the robot is in a walking state. We define two leg groups for a tripod gait \u2014 group 0 with legs [0,3,4]; and group 1 with legs [1,2,5] (see Fig. 5b for leg numbering). For the tripod gait the swing and the stance phase will alternate for each leg group. The Cruse pattern (see Fig. 5a) implements the first three Cruse rules [21] in the following way. The swing phase of one leg inhibits the start of the swing phase of the next leg. The start of the stance phase excites the start of the swing phase of the next leg. The position of the previous leg excites the start of the stance phase. This pattern uses the states of the legs to evaluate the rules and switch between phases. The CPG pattern (see Fig. 5b) implements tripod walking and synchronizes the activity of both leg groups. This pattern uses the reflections of the underlying swing and stance behaviours of both groups to wait for each other to be finished with the swing or stance phases. The CPG Local pattern (see Fig. 5c) also implements tripod walking but with no outer synchronization mechanism. This pattern does not consider in which state the other legs are, and controls each single leg independently. At the beginning a leg group will be motivated to swing and the other to stance, and after that, each leg will take into account it\u2019s own swing and stance reflection to determine if it will swing or stance next. Since, no information from the other legs is used, after some time the generated pattern will be asynchronous and the generated gait is unstable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000297_j.ijleo.2019.164128-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000297_j.ijleo.2019.164128-Figure1-1.png",
        "caption": "Figure 1: Meshed model of maraging steel for analysis in ANSYS APDL.",
        "texts": [
            " \ud835\udc470 = 303 \ud835\udc3e \ud835\udc470 is ambient temperature. The building platform is a mixture of powder and solidified metal. The thermal conductivity of loosely spaced powder in terms of solid is given by \ud835\udc58\ud835\udc5d = \ud835\udc58\ud835\udc60(1 \u2212 \u00d8) (4) Where \u00d8 is the porosity of loosely spaced powder and can be obtained from the following equation. Considering the proportion of powder (density\ud835\udf0c\ud835\udc5d) and solidified metal (density\ud835\udf0c\ud835\udc60)[15]. \u00d8 = \ud835\udf0c\ud835\udc60\u2212\ud835\udf0c\ud835\udc5d \ud835\udf0c\ud835\udc60 (5) In the current study, a prismatic model of size 1500mm x 400 mm x 350 mm is chosen with a mesh size of 5\u00b5m as demonstrated in Figure 1. The whole model is divided into two sections. The lower portion is assigned the properties of solid geometry while the upper part is designated the properties of powder particles in order to replicate the real environment. The symmetry of the model along the Xaxis is utilized in order to reduce the size of the model and computational time. The study is conducted using 4 layers and element birth and death technique is used. A dwell time of 1 second is practiced between the deposition of the consecutive layers, thereby, allowing some cooling between the successive layers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002733_j.addma.2021.101900-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002733_j.addma.2021.101900-Figure16-1.png",
        "caption": "Fig. 16. General specimens dimensions and printing path for each layer.",
        "texts": [
            " For these simulations the distance between deposited filaments was fixed as d = 0.325 mm. The printer settings and additional printing parameters used in the simulations can be seen in Table 1. It should be noted that when E was changed, extrusion speed of material exiting the nozzle also changed proportionally. This speed change is automatically implemented by the printer when it interprets the value of E set in GCode in order to control the feed of material into the hot end. The printing path for each layer is depicted in Fig. 16. It can be observed that unidirectional printed parts were simulated for these tests of VOLCO-X. This was because the contact between deposited filaments plays a major role on the material distribution for these 3D printed parts. In this manner, it was possible to better evaluate the performance of VOLCO-X with the new proposed model for contact between filaments. It is possible to estimate the cross sectional dimensions of specimens, based on the data given in Table 1. With several distances d between printed filaments, 22 as number of filaments per layer, 0.3mm as layer thickness and 4 as number of layers, nominal values for dimensions in Fig. 16 can be estimated: W in the range from 7.150 to 7.975mm and R.Q. Macedo et al. Additive Manufacturing 40 (2021) 101900 t = 1.2 mm. However, such dimensions vary according to actual deposited amounts of material, which vary with the aforementioned parameters Vf and E, as will be seen in the results section. Experiments were performed to validate the simulation results. Specimens of ABS-MG94 (filament manufactured by Faz3D) were produced by FFF in an enclosed 3D printer (GTMax3D Core A2 \u00ae) following the same printing parameters in Table 1 used for the VOLCO-X simulations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000204_j.autcon.2019.102996-Figure18-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000204_j.autcon.2019.102996-Figure18-1.png",
        "caption": "Fig. 18. Subpart B with (a) a quarter of the surface highlighted, and then (b) united to the other geometries.",
        "texts": [
            " The outline of the notch may be modified according to the profile available. Next, the curve is extruded to form the surface of subpart A (Fig. 17b). The hub axis vector guides the extrusion of the subparts related to the corresponding hub. This step completes after top and bottom surfaces cap the notch geometry (Fig. 17c). 3.3.8.2. Generation of subpart B. Subpart B joins a rectangular section from subpart A and a circular section from subpart C. We simplified the circular and rectangular sections into segments of curves to avoid geometry errors in the program. Fig. 18a shows this simplification and subsequent surface generation. All quarters join into one surface; then, the joined surface connects to subparts A and C (Fig. 18b). 3.3.8.3. Generation of subpart C. The segments of mesh lines depicted in Fig. 12a are base for generation of subparts C, which are pipes with predefined radiuses. The start point of the pipe is the intersection of the segments of mesh lines and the external cylinder (Fig. 16). External and internal cylinders define a transition zone from subpart A to subpart C. The part between both cylinders is called subpart B. The radius of the internal cylinder is the same as the radius of the hub, and the radius of the external cylinder is larger than the internal one by any amount necessary for the transition"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000300_bibe.2019.00122-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000300_bibe.2019.00122-Figure14-1.png",
        "caption": "Fig. 14. Structure of a dielectric loaded monopole antenna with optimization",
        "texts": [
            " When is further increased, the value of the objective function jumps over the local minima trap, . This demonstrates the fact the proposed algorithm with large and effective optimization updates has a better chance of avoiding local minima traps than classical quasi-Newton which uses smaller optimization updates. This feature increases the robustness of our proposed algorithm compared to classical quasi-Newton method. The third example is a dielectric resonator loaded antenna example [35]\u2013[37]as shown in Fig. 14. The monopole antenna is loaded with dielectric resonator discs to increase the operating range of the monopole. The height of the monopole from the ground plane is fixed at 8.75 mm and therefore limits the lower end of the operating frequency range of the antenna. Height and height are the heights of the dielectric resonator loaded disc on the monopole. The dielectric resonators have inner radius and outer radii , for each disc respectively. The ground plane size is chosen to be 35 mm 35 mm. The design space vector for the example are chosen based on the sensitivity information"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002766_j.apt.2021.02.019-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002766_j.apt.2021.02.019-Figure2-1.png",
        "caption": "Fig. 2. Schematic illustration of the particle spreading step. The initial packing height is 0.038 cm in the delivery chamber. The lengths and widths of the delivery and fabrication chambers are the same at 0.05 by 0.05 cm each. The layer heights in the fabrication chamber are between 100 and 350 lm.",
        "texts": [
            " Parameter Value Particle mean diameter, l (lm) 30 Particle density, q (kg/m3) 2500 PSD width, r/l (-) 0.3 Poisson ratio, m (-) 10%, 30%, 50%, and 70% Young\u2019s modulus, Y (GPa) 0.01 Hamaker constant, Ha (J) 6:5 10 20 Static friction coefficient (-) 0.3 Rolling friction coefficient (-) 0.002 Restitution coefficient (-) 0.95 Time-step, Dt (s) 5 10 7 Fig. 3. Particle piles: (a) experimental [27], and (b) simulation. V is t 30 lm. Fig. 1 presents frequency versus particle diameter of the number-based lognormal PSDs in this work. Fig. 2 illustrates the schematic diagram of the particle spreading step in particle-based AM [1]. Firstly, the particle bed in the delivery chamber with length and width of 0:05 cm and 0:05 cmis elevated, while the fabrication chamber with length and width of 0:05 cm and 0:05 cm is lowered. Secondly, the blade or roller pushes the particles from the delivery chamber to the fabrication chamber. Lateral displacement and longitudinal compression are generated by the translational and rotational motions of the blade or roller"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001152_ab3f56-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001152_ab3f56-Figure1-1.png",
        "caption": "Figure 1. Configuration of active compression sleeve using wirefabric mechanism with soft pressure sensor.",
        "texts": [
            " The principle of the mechanism is explained here by presenting the manufacturing process and the principles of operation of the overall structure as well as the individual components. The effectiveness of the proposed mechanism was validated by testing the maximum and the minimum variable compression force levels, the variable performance of the desired compression forces, and the controllability of the constant compression force under internal volume changes. The compression sleeve based on the wire-fabric mechanism mainly consists of a two-layered (i.e. outer and inner) fabric structure, as shown in figure 1. The outer layer was used to generate the compression force by pulling both ends with a wire actuation system to generate the compression force. A soft pressure sensor was integrated between the outer and inner layers to monitor the compression force generated by the outer layer in real time. It is then possible to estimate the compression force applied to the body using the signal from the soft sensor. The pulling force of the wire pulling is controlled with the sensor feedback to keep the compression in a desired level. The outer and inner layers of the sleeve were fabricated as shown in figure 1. The inner layer was worn on the body like a sleeve, and it supports the main components, such as the actuation system and the soft pressure sensor. It also prevents any possible damage to the skin during wire driving. In consideration of these requirements, spun-silk fabric was used for the inner layer material, as it is soft and elastic. Elastic bands are used for padding on both ends of the inner shell for stable attachment of the device to the targeted body part during wearing. The outer layer generates the compression force to the body by pulling the ends of the sleeve together by wires"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001096_s00170-019-04076-4-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001096_s00170-019-04076-4-Figure1-1.png",
        "caption": "Fig. 1 Mounting of worm and grinding wheel",
        "texts": [
            " The worm tooth surface is strictly composed of Niemann tooth profile, non-generated tooth surface, and fillet surface with circular arc. The mathematical model of Niemann profile is formulated in this chapter because Niemann profile in these tooth surfaces has the majority and has an influence on the meshing performance directly. The worm tooth surface with Niemann profile is usually generated as an enveloping about a series of the grinding wheel with convex circular arc. The tooth surface of worm wheel is considered as the conjugate tooth surface which is generated as an enveloping about a series of the worm tooth surface. Figure 1 shows the mounting of the worm and grinding wheel. The crossing angle between the axes of the grinding wheel and worm is equal to the lead angle \u03b3 of the worm. Figure 2 shows the coordinate system Og-xgygzg of the grinding wheel and the profile of grinding wheel with convex circular arc in cross section xg = 0. P is the arbitrary point set on convex circular arc of grinding wheel. u is a variable parameter which represents the position on the curved line of the grinding wheel and means the depth of cut"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001993_s00170-020-06278-7-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001993_s00170-020-06278-7-Figure12-1.png",
        "caption": "Fig. 12 a\u2013d Grinding parameters of experimental",
        "texts": [
            "2, and the wheel position and Table 2 Tool flute and grinding wheel parameters Cutting tool parameters Two-pass parameters Wheel parameter No. Tool radius R (mm) Inner core radiusRc (mm) Rake angle \u03b3o (\u00b0) Edge width Ew (mm) Outer core radius Rcb (mm) Joint radius rj (mm) Flute angle ratio \u03b4 (%) Wheel width hW (mm) A 3 1.9 6 0.8 2.7 2.428 64 6.1 B 4 2.56 9 1.2 3.6 3.246 orientation are installed as Table 3. The cross-sections of ground helical grooves produced according to the proposed method were also examined using a KEYENCE VHX-5000 digital microscope. Figure 12 a and b show the cross-sections after the first and the second grinding passes for group A. Figure 12 c and d depict the first and second grinding passes for group B. To make it easier to assess the grinding parameters results intuitively, the annotations in Fig. 12 show the measured values. And, the value of grinding parameter results is measured by the ZOLLER ThreadCheck machine. To be able to further validate the smoothness of the grinding marks between the two passes, the ground flutes of the two tools were also observed by using a KEYENCE VHX-5000 digital microscope. Figure 13 a and b show the two tools, A and B, respectively. There are no obvious grinding marks on either of them. The ground parameters of the two tools, inner core radius Rc, rake angle \u03b3o, edge width Ew and outer core radius Rcb (corresponding to Fig. 12 b and d), were measured using a ZOLLER ThreadCheck machine. The results are listed in Table 5. The error ratio is calculated by subtracting the design value from the measured value and dividing by the design value. To minimize any possible errors, the measurements were performed three times. The maximum average errors for the inner core radius, rake angle, edge width, and outer core radius of the two ground tools were 1.1%, 2.0%, 3.4%, Table 4 Simulation results No. Results/error ratio Inner core radius Rc (mm) Rake angle \u03b3o (\u00b0) Edge width Ew (mm) Outer core radius Rcb (mm) A 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000907_j.talanta.2019.02.093-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000907_j.talanta.2019.02.093-Figure1-1.png",
        "caption": "Fig. 1. Scheme of the setup for dynamic gas extraction.",
        "texts": [
            " Halogen working solutions were prepared by diluting the initial solution immediately before use. A stock solution containing 0.01 mg mL\u22121 of halide ions was prepared by dissolving the weighed portions of the corresponded salts (NaCl, KBr) in 500 mL of water. The precise concentration was established by means of the argentometric titration method. Working solutions were prepared by diluting the stock solution immediately before use. Dynamic gas extraction was carried out using setup represented in Fig. 1. It included a glass vessel for analyzed solution (1), blanked off with a rubber stopper (2), a holder of test strips (3) with a test strip grasped, an air microcompressor (4) connected via a polymer hose (5) with a glass bubbler (6) sealed in the vessel. To make circulate air through the system an air microcompressor \"Hailea Aco-6601\" and an autonomic air microcompressor \"Jebo 3800DC\" were used in laboratory and in on-site conditions respectively. When carrying out determinations in chlorinated water on-site, the vessel (element 7 on Fig. 1) containing MO solution (\u0421=0.1 g L\u22121) acidified by sulfuric acid was placed between the air microcompressor and the reaction vessel. This manipulation is required to purify air from chlorine contamination. Diffuse reflectance spectra of the reaction zone of the paper-based sensor (PBS), as well as absorbance of solutions, were recorded on a Varian Cary 5000 UV\u2013Vis\u2013NIR (Agilent) spectrophotometer equipped with diffuse reflectance accessories DRA \u2013 2500 or a standard cell holder. The \u0440\u041d of solutions was controlled by a pH-meter \u0440\u041d-150 MI (OOO \"Measurement technology\")",
            " Therefore, the indication should be provided either via coloration of the solution arising after contacting with PBS or via carrying out the reaction out of the analyzed solution and using bleaching degree of the pink color of the PBS as a quantitative characteristic of the method. In the first case, the technique will be characterized by low selectivity due to interfering of other oxidants being able to bleach MO. Hence, conducting the reaction on the PBS surface out of the analyzed solution is more favored. In order to extract halogens from the sample, we have accomplished gas extraction by the help of the setup presented in Fig. 1. During longtime contact of the sensor with air flow (more than 45 min) desorption of the dye is not observed. This fact is confirmed by the constant value of the reflectance diffusion coefficient. The technique consists of the following stages: halogen gas extraction, the interaction of the halogen with the dye attached to the solid support, and determination of the coloration intensity of the reaction zone of PBS that changes depending on the concentration of halogen in solution (Fig. 4). Conducting the reaction out of the sample provides selectivity of the analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001019_rpj-07-2018-0182-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001019_rpj-07-2018-0182-Figure12-1.png",
        "caption": "Figure 12 Explicit analysis contact pressure distribution",
        "texts": [
            " Field mapping option available in FEA software is used to transfer the contact pressure distribution from the explicit analysis to a static analysis where only the bottom die is considered for analysis, while other boundary conditions remain same. Contact pressure obtained during the forming is considered as a static load and applied to die in the static analysis. Figure 11 shows the mapped surfaces, boundary condition and load because ofmapping. Contact stress distribution plot for the explicit analysis is shown in Figure 12. Maximum contact stress is observed where the Metal bellow hydroforming Prithvirajan R. et al. Rapid Prototyping Journal D ow nl oa de d by N ot tin gh am T re nt U ni ve rs ity A t 0 2: 47 3 1 M ay 2 01 9 (P T ) bellow root region will be formed and reduces towards the crest forming region and lesser at other regions where there is no contact. The maximum pressure 31 MPa acts outward in the radial direction to the die which tends to compress the die material. This radial contact stress is because of the interaction of tube external surface with the die as the tube expands as a result of applied forming pressure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003774_s0020-7403(96)00063-x-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003774_s0020-7403(96)00063-x-Figure1-1.png",
        "caption": "Fig. 1. The two phases kinematical model of the axisymmetric folding of circular tubes.",
        "texts": [
            " The analysis that was briefly outlined is now developed in more detail. 575 576 M. Avalle and G. Belingardi The tubes considered here have an outer diameter of 100 mm, a wall thickness of 2 mm and an axial length of 400 mm. All of them are aluminium, made by cold extrusion. All of them were submitted to annealing thermal treatment in order to recover the ductility typical of aluminium. At the end of the annealing treatment, the material has an elastic limit strength of about 80 MPa and a plastic strain-hardening modulus of about 2000 MPa. Figure 1 shows the kinematics of the collapse during the formation of one fold, according to the detailed model developed in a previous paper [6, 7]. Each fold consists of two conical surfaces, caused by the radial displacement of a portion of the tube wall. The axial length of the portion of tube involved in the formation of each fold is termed fold length. It is clear that the conical surfaces are the result of the large plastic bi-dimensional bending of a cylindrical shell. The material is submitted to large plastic strain both where plastic hinges take place and where stretching takes place as a consequence of the increase of the radius",
            " Strains can be elaborated as a function both of time and space giving a clear description both of the strain field and of its evolution from the first buckling phase which occurs after yielding (already in the plastic range) to the folding phase. 4. R E S U L T S 4.1. Time-histories of strains Figures 4 and 5 show the diagram of the measured strains as function of the displacement of the testing machine cross-head which means the depth of the progressive collapse. In the same figures the diagram of the crushing force is also shown, so that it is easy to see the connection of the progressive folding (the formation of each fold generates two peaks of the crushing force as explained by the simplified model drawn in Fig. 1) with the strain signals. At the first raising of the crushing force there is a uniform negative (compression) value of the axial strains and a uniform positive (tensile) value of the circumferential strains. The material flow strength has been reached and the following buckling, that determines the first fold, develops when the material is already in its plastic phase. Figure 6 shows the stress-strain diagram that can be obtained by calculating the ratio of the axial force to the area of the tube cross-section (far from the folding zone this is a purely axial stress) and by relating this with the axial strain measured by one of the strain gages",
            " It is also worth noting that the circumferential strains measured by strain gages 15 (placed at the same axial coordinate of the strain gage 0) and 24 (placed at the same axial coordinate of the strain gage 9) initially decrease until they become negative and then increase to positive high values. This can only be explained by an initial inward movement of the original cylindrical surface of the tube during the fold initialization phase, then followed by the outward motion of the surface during the fold formation phase, as illustrated in Fig. 1 I-6, 7]. Figure 8 shows the diagram of the circumferential strains as a function of the crush displacement, calculated on the basis of the kinematical model described in Fig. 1. The diagram is drawn on the basis of the experimental value of the fold length (2h = 27 ram) and of the eccentricity (y/h = 0.645). This diagram compares well with that of Fig. 5 concerning the shape of the strain curves. Large differences are visible between the strain magnitudes. But it is necessary to consider two facts: (1) The experimental maximum value of the radial displacement is about 5.5 mm while in the calculation y was equal to 8.7 mm and therefore 60% larger. Since the circumferential strains are proport ional to the radial displacement, the theoretical values are 60% larger than the measured ones. If such a reduction is performed the strain values are comparable. (2) Although the strain gage size chosen was rather small, it gives an averaged value of the strain over the area where it was glued. Thus strain measures are valid over a small area but cannot be related to one point. In the present application the strain gage size cannot be considered very small with respect to the fold length, thus peak values are decreased to some extent by averaging. A folding kinematic, similar to Fig. 1, is schematically depicted in Fig. 9, taken from a recent paper by Wierzbicki and co-workers [10]. The obtained strain measurements validate the basic assumptions of the kinematical theoretical model proposed in our papers [6, 7] as well as in Ref. [10]. The same conclusions have been drawn in another recent paper [11] on the basis of measures of the fold geometry performed on some folded tube sections. It is not possible to draw a diagram of the axial strains calculated on the basis of the kinematic model described in Fig. 1, because the values are indefinitely large as the plastic hinges and zero elsewhere, and therefore do not compare with those of Fig. 4. Further comments are developed at the end of the next section. 4.2. Strain field The strains, previously presented in their evolution with time, can now be examined from a spatial point of view. In Fig. 10(a) and (b) the axial strains are plotted as a function of the axial position of the strain gage, each line corresponds to a different value of the testing machine cross-head displacement",
            " The folding length measured by the distance of two subsequent strain peaks after folding is about 27 mm and is quite different from the elastic wavelength, while is in reasonable agreement with the results of the theoretical models by Alexander [2], Abramowicz and Jones [3] and Singace and co-workers [11]. It was already noted that the fold radial span (i.e. the difference between external and internal radius) is rather less than half the fold length. Moreover the shape of the strain curves of Fig. 10(b) are quite different from the peaked diagram required by the kinematical model of Fig. 1. Therefore the fixed plastic hinges model [2, 6, 7, 11] is not adequate to describe the wall deflection in the meridian section and a moving plastic hinges model (Fig. 9) might appear better. But this conclusion is not supported by the axial strain diagram of Fig. 10(b). In fact, if the shell axial curvature can be assumed proportional to the axial strain, it is not constant, as it should be according to the assumption made by Wierzbicki and co-workers 1-4, 10]. 5. C O N C L U S I O N S The crushing behavior of a circular tube submitted to quasi-static axial loading has been considered by experimental analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001024_j.ymssp.2019.05.021-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001024_j.ymssp.2019.05.021-Figure12-1.png",
        "caption": "Fig. 12. Calibration setup.",
        "texts": [
            " (58)), but in this case it is not possible to estimate f and sc simultaneously from the ratios or differences of successive amplitudes. In practice, preliminary experiments can be conducted to identify whether a viscous or dry friction model best fits the system, and then use one or the other. The torsion spring constant ks introduced in Eq. (48) was identified through a linear regression between applied torques and observed angular displacements. To find ks the torsion platform was fixed sideways and a weight was hanged from a symmetrical bar attached to the shaft to apply a known torque around the axis of rotation (Fig. 12). The spring was unloaded by loosening the shaft attached to the bottom of the spring and letting it unwind. The encoders were zeroed in the unloaded position, with the weight hanging directly below the axis of rotation. The calibration dataset was generated by applying a torque to the shaft attached to the bottom of the spring and measuring the resulting rotation h. The angular displacements measured by both encoders were recorded. The spring\u2019s net angular displacement, h t\u00f0 \u00de, is calculated from the measurements of the top and bottom encoders, h1 t\u00f0 \u00de and h2 t\u00f0 \u00de as h t\u00f0 \u00de \u00bc h2 t\u00f0 \u00de h1 t\u00f0 \u00de \u00f075\u00de The torque applied to the shaft, ss t\u00f0 \u00de, is calculated by ss t\u00f0 \u00de \u00bc mclg sin h1 t\u00f0 \u00de \u00f076\u00de where mc is the mass used for calibration (0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001999_ecce44975.2020.9235600-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001999_ecce44975.2020.9235600-Figure13-1.png",
        "caption": "Fig. 13. Magnetic flux distributions of the proposed VFMM and the target IPMSM under no-load condition at 17000 rpm.",
        "texts": [
            " Fig. 12 shows no-load line-to-line voltages in T1 and T3 periods. Difference between the fundamental waveform amplitude of line-to-line voltage before and after applying maximum load current is very slight of 0.96%. Hence, it is obvious that the proposed VFMM can avoid demagnetizing VPMs unintentionally. In the proposed VFMM, first layer CPMs contribute to enhancing durability against demagnetization of VPMs. As a result, unintentional demagnetization as a major problem of VFMMs can be prevented. Fig. 13 shows magnetic flux distributions of the proposed VFMM in different magnetization ratio of VPMs and the target IPMSM. Compared with case of the magnetization ratio of 0%, the magnetic flux density of the stator is obviously lower when the magnetization ratio is -100% as shown in Fig. 13(a) and (b). This is because VPMs draws magnetic flux from CPMs, and resultantly, the magnetic short-circuit is composed in the rotor. Fig. 13(c) indicates the target IPMSM, and its magnetic flux distribution in the stator core is almost similar to the proposed VFMM having VPMs whose magnetization ratio is 0%. For the above results, it has been found that the proposed VFMM can achieve almost same and lower air gap magnetic flux density compared with the target IPMSM at the same time. In other words, the Fig. 12. Comparison of no-load line-to-line voltage in T1 and T3 periods. -150 -100 -50 0 50 100 150 0 0.001 0.002 0.003 0.004 0.005 Li ne -t o- lin e vo lta ge [V ] Time [s] T1 125"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002765_s11071-021-06327-0-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002765_s11071-021-06327-0-Figure8-1.png",
        "caption": "Fig. 8 Track pitch and sprocket pitch",
        "texts": [
            " Depending on track pitches, there are three cases when a track meshes with a sprocket: equal-pitch meshing, sub-pitch meshing and extra-pitch meshing. In this section, the established contact model is utilized to analyze these three meshing conditions combined with MBD simulation. A fourth-order Runge\u2013Kutta integration method is selected with constant step size as h = 1 9 10-4 s and the minimum penetration determination e = 1 9 10-5 m. For each case, the simulation parameters are the same except for rp and ra. Besides Table 1, the remaining simulation data are listed in Table 2. 3.1 Equal-pitch meshing As demonstrated in Fig. 8, the inner circle center distance of adjacent track pins is defined as a track pitch PT. The distance between adjacent teeth of the meshing positions on the pitch circle is a sprocket pitch Pz, which is evaluated as PZ \u00bc 2\u00f0rp \u00fe r5 \u00fe D45\u00de sin\u00f0p=Z\u00de \u00f032\u00de where Z depicts the number of teeth. In this case, PT = PZ and rp = 0.256 m, ra = 0.319 m. The normal contact forces between a track pin with three surfaces of an arbitrary tooth groove are displayed in Fig. 9. In the simulation, a tracked vehicle is traveling on a flat road at a constant speed of 1 m/s"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000455_1.4032471-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000455_1.4032471-Figure3-1.png",
        "caption": "Fig. 3 Sketch and coordinate systems for the applied CNC machine",
        "texts": [
            " (13)\u2013(18) which can be used to determine the pinion machine settings for the applied CNC machine. 5.2 Mathematical Model of the Applied Multi-Axis CNC Machine. The proposed global synthesis approach can be implemented by employing a free-form five-axis CNC machine. Fortunately, more and more five-axis machine tools are adopted for the manufacture of spiral bevel and hypoid gears in the industrial field. The sketch and employed Cartesian systems of the applied five-axis CNC machine are presented in Fig. 3. As shown in Fig. 3, S(c){Oc; e \u00f0c\u00de 1 , e \u00f0c\u00de 2 , e \u00f0c\u00de 3 } is the reference sys- tem of R(c); S(m){Om; e \u00f0m\u00de 1 , e \u00f0m\u00de 2 , e \u00f0m\u00de 3 } is the reference system of machine. S(w){Ow; e \u00f0w\u00de 1 , e \u00f0w\u00de 2 , e \u00f0w\u00de 3 } is the reference system of R(1), and Ow is chosen at the crossing (intersecting) point of the gear axis and the pinion axis. X, Y, Z, c, w and d denote the applied CNC machine settings which can be represented as X \u00bc ax0 \u00fe ax1/\u00fe ax2/ 2 \u00fe ax3/ 3 \u00fe ax4/ 4 \u00fe ax5/ 5 \u00fe ax6/ 6 (19) Y \u00bc ay0 \u00fe ay1/\u00fe ay2/ 2 \u00fe ay3/ 3 \u00fe ay4/ 4 \u00fe ay5/ 5 \u00fe ay6/ 6 (20) Z \u00bc az0 \u00fe az1/\u00fe az2/ 2 \u00fe az3/ 3 \u00fe az4/ 4 \u00fe az5/ 5 \u00fe az6/ 6 (21) c \u00bc ac0 \u00fe ac1/\u00fe ac2/ 2 \u00fe ac3/ 3 \u00fe ac4/ 4 \u00fe ac5/ 5 \u00fe ac6/ 6 (22) Journal of Mechanical Design MARCH 2016, Vol",
            "org/ on 02/10/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use w \u00bc w \u00fe / (23) where axi, ayi, azi, and aci (I\u00bc 0 to 6) are real coefficients; / is the parameter of generation and w is the initial value of w. For the given cradle machine-tool settings, the equivalent settings for applied CNC machine can be easily obtained by transformation [35]. 5.3 Determination of the Relative Position and Motion Between the Tool and the Workpiece. Now, let us consider the generation of R(1) using the CNC machine shown in Fig. 3. By employing the references systems and the component notation presented in Ref. [38], the tensor equations (13)\u2013(18) can be transformed into the equivalent scalar equations. In this paper, we will consider a straight-line profile for the blades of tool. (z(cL), r(cL), h(cL)) indicates the point on the tool surface corresponding to M \u00f01\u00de i . Therefore, z(cL) and r(cL) are linearly related. For the specific M \u00f01\u00de i and the given r(cL), the machine settings X, Y, Z, c, w, and h(cL) can be determined according to Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001207_rpj-07-2018-0171-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001207_rpj-07-2018-0171-Figure3-1.png",
        "caption": "Figure 3 (a) Design space; (b) 1st design stage; (c) 3rd design stage; and (d) 6th design stage (manufactured).",
        "texts": [
            " The objective function was the total weight of the axle carrier and because software used the SIMPmethod, the variable was density r . The optimization also needs to satisfy the elastic equilibrium for global stiffness matrix K andmatrix of loads f in every point u (4): min r : C (1) Subject to: V xVC (2) t tmin (3) And satisfy the equilibrium constraint: Ku \u00bc f (4) The main criteria for designing the envelope part were the available space in the wheel rim and preventing collision with suspension rods [Figure 3(a)]. In the first optimization run, the calculation parameters were set on 25 per cent of the used material and the minimal wall thickness of 7mm. The weight of the axle carrier in the first design stage was 875g. Additional restricted areas were added with respect to functionality, for example, the free space for the nut andwrench formounting [Figure 3(b)]. The second design stage was set on 30 per cent and theminimal wall thickness of 5.5mm.Theweightwas reduced to 596g. The specific behaviour of solidThinking Inspire (described earlier) was used in the third design stage. As the material is more functionally distributed near the forces, the loading of the axle carrier was applied as in FEM analysis (Figure 1), and the material distribution in the central part of the carrier was optimized. In this design stage, the supporting beam was added to prevent a flange fromdeformation [Figure 3(c)]. The basic shape of the axle carrier was then upgraded during three follow-up design stages. The component orientation on the building platform was chosen and the shape was additionally optimized with regard to additive manufacturing Topologically optimized axle carrier Ond rej Vaverka, Daniel Koutny and David Palousek Rapid Prototyping Journal Volume 25 \u00b7 Number 9 \u00b7 2019 \u00b7 1545\u20131551 and expected machining. Areas that should be milled (under bolts) were elevated for easier machining [Figure 4(a)]",
            " To improve the mesh quality, a mapped mesh was used for rectangular or circular surfaces (holes and mounting surfaces for bearings). The final mesh contained approximately 205,000 elements and 386,000 nodes, whereas 80 per cent of the elements had element quality better than 0.5. In terms of degrees of freedom (DOFs), the model size was approximately 1.15 106 with threeDOFs per node. A direct solver, time step of 0.1 s, unsymmetrical Newton\u2013Rhampson method, constant stabilization, weak springs and 5 per cent convergence on the total deformation of the axle carrier were set for the analysis. The final CAD model [Figure 3(d)] was added with machining allowances of minimum 1mm [Figure 6(a)] and exported in the STL format. The polygonal mesh was then finally smoothed in the Autodesk Meshmixer (Autodesk, Inc., San Rafael, CA, USA) to eliminate remaining sharp edges and minimize the stress concentrators caused by defects of the mesh [Figure 6(b)]. Data for additive manufacturing were prepared in software Materialise Magics 21.11 (Materialise, Leuven, Belgium). The axle carrier was placed on a building platform with nearly minimal height to avoid residual stress and further deformation in the central area used for mounting the bearings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001460_s00170-019-04874-w-Figure21-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001460_s00170-019-04874-w-Figure21-1.png",
        "caption": "Fig. 21 Assembly model of the WBCFN. a CAD model. b Physical object",
        "texts": [
            " Those are also consistent with the results described in Fig. 17. Combining with analysis on the objectives, the design scheme P3 (L = 100 mm, a = 25\u00b0, w = 1.5 mm) is selected as the final optimal solution, comprehensively. The experimental validation for powder flow and direct laser deposition is carried out to ensure the reliability of this model and calculation. The WBCFN used in the experimental validation is the final optimal scheme discussed. The whole assembly model of the WBCFN is shown in Fig. 21. Figure 22 shows the powder flow comparison between detected image and simulated concentration profile. As can be seen, the results of simulation and experiment reveal similar distribution of the powder concentrations both in transverse and longitudinal slices. The relationship between the powder image gray value and powder concentration is given in Eq. (10). Several images are taken to calibrate the proportional constant. The calibration process is as follows: First, a local 5 \u00d7 5 mm region is defined in the center of focal point"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001905_j.mechmachtheory.2020.104095-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001905_j.mechmachtheory.2020.104095-Figure10-1.png",
        "caption": "Fig. 10. Constraint relationship between central chain and kinematic chains: (a) Plane y B O B z B (b) An axonometric drawing to show the constraint relationship on plane x B O B z B . In Fig. 10 (a), point P 2 , P 3 , a, b, c, d, O p \u2019 and angle \u03b11, \u03b81 x, \u03b1 are on plane y B O B z B . In Fig. 10 (b), point P 1 , a, b, c, d, P c , O p \u2019 and angle \u03b21, \u03b81 y, \u03b2 are on plane x B O B z B .",
        "texts": [
            " 5 l 4 c cos ( \u03b81 + \u03b82 ) (16) Point P 0 and O p \u2019 can be expressed, respectively: { B \u2032 x P 0 = ( l 2 c + l Mc + l 3 c ) sin \u03b81 + l 4 c sin ( \u03b81 + \u03b82 ) B \u2032 z P 0 = ( l 2 c + l Mc + l 3 c ) cos \u03b81 + l 4 c cos ( \u03b81 + \u03b82 ) (17){ B \u2032 x O \u2032 P = L c sin \u03b81 B \u2032 z O \u2032 P = l 1 c + L c cos \u03b81 (18) where, L c = l 2 c + 0 . 45 l Mc . Then, according to Fig. 9 , we can get Eqs. (19) and (20) B \u2032 x 2 O \u2032 P = L 2 c sin 2 \u03b81 = X 2 + Y 2 (19){ Y = L c sin \u03b81 x X = L c sin \u03b81 y (20) 2.4. Constraint relationship between central chain and kinematic chains In this section, we give the constraint relationship between the pose of the moving platform and the generalized coordinates of kinematic chains. In Fig. 10 , we project the deformation of the central chain and the cantilever beam, the pose of moving platform to plane y B O B z B and plane x B O B z B , respectively. To facilitate showing the constraint relationship on plane x B O B z B , we give an axonometric drawing in Fig. 10 (b). In Fig. 10 (a), triangle aO p \u2019b is similar to triangle bP 3 c . We get \u03b11 = aO p \u2019b . Line bc is parallel to axis z B . That makes \u03b81 x = aO p \u2019b . Then, considering that \u03b1 = aO p \u2019b , we can get the relationship shown in Eq. (21) . \u03b81 x = \u03b11 = \u03b1 (21) Next, considering that triangle P 2 dP 3 is a right triangle, we can get the relationship shown in Eq. (22) . sin \u03b11 = P 2 d P 2 P 3 = z p2 \u2212 z p3 \u221a 3 L (22) where, the meaning of L is shown in Fig. 7 (b). In Fig. 10 (b), P c is the center of P 0 P 2 . That makes z pc = 0.5( z p 2 + z p 3 ). We can find that triangle aO p \u2019b is similar to triangle bP 3 c , line bc is parallel to axis z B and \u03b1 = aO p \u2019 b . Then, we can get Eq. (23) \u03b81 y = \u03b21 = \u03b2 (23) Triangle P 1 dP c is a right triangle, that means we can get Eq. (24) . sin \u03b21 = P 1 d P 1 P c = z p1 \u2212 0 . 5 ( z p2 + z p3 ) 1 . 5 L (24) Considering that the deformation of the central chain (about 30 \u03bcm) and the cantilever beam (about 0.03 \u03bcm) is not significant, the value of \u03b81 x, \u03b81 y , and \u03b82 can be regarded as insignificant"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003727_1.2831176-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003727_1.2831176-Figure7-1.png",
        "caption": "Fig. 7 Workpiece geometry and cutter paths used in cutting tests",
        "texts": [
            " Machining tests were carried out on two separate set-ups consisting of an experi mental two-slide test bed and a vertical machining center 472 / Vol. 119, NOVEMBER 1997 Transactions of the ASME Downloaded From: http://manufacturingscience.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use (Okuma MC4VAE). The cutting forces were measured with a three-component dynamometer (Kistler Model 9257A) and collected with a Macintosh II based data acquisition system. Cast iron workpieces of varying geometric complexity (Fig. 7) and carbide inserts were used in the experiments. A 102 mm diameter face milling cutter with 15 deg lead angle was used. The two slide test bed was used for the linear cutter feed path experiments and the vertical machining center for the curvilin ear cutter feed path test. Only dry single tooth milling was used in all the experiments. The cutting conditions consisted of a fixed axial depth of cut of 1.27 mm and three levels of feed (0.15 mm/tooth, 0.178 mm/tooth, 0.381 mm/tooth). The regions and positions where cutting forces were collected are also marked in Fig. 7. Prior to using the cutting force model outlined in the sections above, the cutting force coefficients Kc and K, need to be deter mined from the average force data collected during cutting tests. The coefficients, as a function of the average undeformed chip thickness 7C, are estimated as, Kc = 51135.5 7;031345 (N/mm2) (27) K, = 6463.5 7t: a64434 (N/mm2), (28) Using the calibrated cutting force coefficients (Kc and K,) database, the three cutting force components (X, Y and Zforces) were predicted for machining the workpieces with vari able geometry",
            " It should be pointed out that the circular holes in the workpiece shown in Fig. lb are for fixturing pur poses. Forces shown (in Fig. 8 and Table 1) are for the solid portion of the workpiece as indicated in Fig. lb. It is seen that the predicted force profiles are in good agreement in both shape and magnitude with the corresponding measured forces. From Table 1, it can also be seen that there is a fairly good match between the model predictions and the observed cutting forces for machining a slotted workpiece (Fig. 7a) with the cutter fed in the \u2014X direction (Case #4). A comparison of the measured and predicted cutting forces for this case are shown in Fig. 9. It is seen that the cutting force profile over the duration of one cutter revolution consists of three segments. The smaller segments contain the forces generated during the cutting of the (side) ribs, while the middle (large) segment is due to the cutting of the (middle) solid workpiece. Figure 9 shows that the model captures the observed variations in the force profile reasonably well"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000567_j.jsv.2016.04.020-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000567_j.jsv.2016.04.020-Figure3-1.png",
        "caption": "Fig. 3. (a) Four subsystems coupled via a planar wedge in the main transmission (b) Subsystems coupled in the two lateral directions.",
        "texts": [
            " A complex stiffness kjn\u00fe i\u03c9cjn is used when viscous damping is considered, where, cjn is the normal damping, i and\u03c9 are the imaginary unit and angular frequency; kjx, kjrz are the lateral stiffness and anti-yawing stiffness in XY plane, which are also resultant stiffness come from the rolling ball's contact compliances; kjz, kjrx are the lateral stiffness and anti-pitching stiffness in YZ plane; f jn; f jx; f jz; Tj rx; Tj rz are the inner forces and torques of the interface. The first equation in Eq. (1) is the new type of joint interface stressed in this paper, which connects four-DOF displacements via a one-DOF joint. Dynamic subsystems are connected (coupled) through the ballscrew-nut interface as illustrated in Fig. 3. They were divided into two groups i.e. the four subsystems coupled via a planar wedge in the main transmission direction (Fig. 3(a)) and the subsystems coupled in the two lateral directions (Fig. 3(b)). The dynamic subsystems considered are: A, B, E, G, the rotational, axial, two bending motion and vibrations of the ballscrew; C, composing subsystems contributing to the nut's axial movement C \u00bc C1\u00feC2\u00feC3; C1, C2, C3, the axial, yawing, pitching motion and vibrations of the slide carriage; D, the rolling motion and vibrations of the slide carriage; F, H are the two lateral motion and vibrations of the slid carriage along X and Z. The dynamics of the two groups, i.e. the dynamics of main transmission and that of the lateral vibrations, are not independent",
            " 1 is summarized [13]: First, the component (subsystem) receptances are h11 \u00bc x1=f 1 , h2a1 \u00bc x2a=f 1 , h12a \u00bc x1=f 2a , h2a2a \u00bc x2a=f 2a , and h2b2b \u00bc x2b=f 2b , and the displacement compatibility condition and the force equilibrium condition of the flexible coupling are \u00f0k\u00fe i\u03c9c\u00de\u00f0x2b x2a\u00de \u00bc f 2a (2) f 2a \u00bc f 2b (3) Then, using the component receptances, we have the coordinate displacements for component I, and II: x1 \u00bc h11f 1\u00feh12af 2a (4) x2a \u00bc h2a1f 1\u00feh2a2af 2a (5) x2b \u00bc h2b2bf 2b (6) Substituting Eqs. (5) and (6) into Eq. (2), and eliminating f 2b using Eq. (3), we have f 2a \u00bc h2a2a\u00feh2b2b\u00fe 1 k\u00fe i\u03c9c 1 h2a1F1 (7) where, f 1 in Eq. (4) is replaced with F1 in the assembly Finally, substituting Eq. (7) into Eq. (4), we have x1 \u00bc h11 h2a2a\u00feh2b2b\u00fe 1 k\u00fe i\u03c9c 1 h2a1 \" # F1 (8) And, the assembly receptance derived is h011 \u00bc X1 F1 \u00bc x1 F1 \u00bc h11 h2a2a\u00feh2b2b\u00fe 1 k\u00fe i\u03c9c 1 h2a1 (9) For the new condition, from the connections in Fig. 3(a) we have the displacement of the subsystems\u2019 interface coordinates: xa1 \u00bc \u03b8Iry ,xb1 \u00bc \u03b4Iry, xc1 \u00bc \u03b4IIry,xd1 \u00bc \u03b8IIry. Transmission ratios from interface coordinates displacements to normal displacements are defined: ina \u00bc ind \u00bc \u00f0d sin \u03b2\u00de 1 , inb \u00bc inc \u00bc cos 1\u03b2. Hence, according to the first equation in Eq. (1), the interface displacement compatibility equations can be written as: \u00f0xI xII\u00de\u00f0kjn\u00fe i\u03c9cjn\u00de \u00bc f jn \u00bc f I xI \u00bc xa1=i n a \u00fexb1=i n b xII \u00bc xc1=i n c \u00fexd1=i n d (10) and the force equilibrium equations is: f I \u00bc f II ina f a1 \u00bc inbf b1 \u00bc f I inc f c1 \u00bc indf d1 \u00bc f II (11) where, xI, xII, fI, and fII are the normal displacements and forces at sides of component I and II; fa1, fb1, fc1, fd1 are the inner forces acted upon the coordinates. Using the subsystem FRFs, the displacements of the coordinates before connection can be expressed as: xa1 \u00bc f a1ha1a1 (12) xa2 \u00bc f a1ha2a1 (13) xb1 \u00bc f b1hb1b1 (14) xc1 \u00bc f c1hc1c1\u00feFc2hc1c2 (15) xd1 \u00bc f d1hd1d1 (16) Fc2 in Eq. (15) is the external force acted upon coordinate c2 as shown in Fig. 3(a). Substituting Eq. (11) into Eq. (13), yields: xa2 \u00bc f I ha2a1 ina (17) Substituting Eqs. (12), (14)\u2013(16) into Eq. (10), and eliminating fa1, fb1, fc1, and fd1 using Eq. (11), we get: Fc2 \u00bc f I ha1a1 \u00f0ina \u00de2 \u00fehb1b1 \u00f0inb\u00de2 \u00fehc1c1 \u00f0inc \u00de2 \u00fehd1d1 \u00f0ind\u00de2 \u00fe 1 kjn\u00fe i\u03c9cjn \" # hc1c2 inc 1 (18) Dividing Eq. (17) by Eq. (18), yields: h0a2c2 \u00bc xa2 Fc2 \u00bc ha2a1 ina ha1a1 \u00f0ina \u00de2 \u00fehb1b1 \u00f0inb\u00de2 \u00fehc1c1 \u00f0inc \u00de2 \u00fehd1d1 \u00f0ind\u00de2 \u00fe 1 kjn\u00fe i\u03c9cjn \" # 1 hc1c2 inc (19) Eq. (19) is the derived prediction equation for h'a2c2, a cross-subsystem displacement FRF of the assembly system",
            " Similarly, the assembly FRFs between other coordinates can be derived, and were arranged as a concise matrix form: h0a2a2 h0a2c2 h0a2c3 h0c2a2 h0c2c2 h0c2c3 h0c3a2 h0c3c2 h0c3c3 2 64 3 75\u00bc ha2a2 0 0 0 hc2c2 hc2c3 0 hc3c2 hc3c3 2 64 3 75 ha2a1 ina hc2c1 inc hc3c1 inc 2 66664 3 77775 ha1a1 \u00f0ina \u00de2 \u00fehb1b1 \u00f0inb\u00de2 \u00fehc1c1 \u00f0inc \u00de2 \u00fehd1d1 \u00f0ind\u00de2 \u00fe 1 kjn\u00fe i\u03c9cjn \" # 1 ha1a2 ina hc1c2 inc hc1c3 inc 2 66664 3 77775 T (20) Eq. (20) is the solution to the receptance coupling problem of the multi-subsystem connected via a wedge mechanism in Fig. 3(a). The receptance coupling Eq. (20) is well structured. The right hand side is a summation of two terms: the subsystem FRFs and the coupling correction. The latter term composes of three multipliers. They are the output response point factor, the coupling effect factor, and the input force point factor, in sequence. The output response point together with the input force point indicates the path of coupling effects. The coupling effect factor indicates the coupling strength and was discussed here"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000441_1.3662591-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000441_1.3662591-Figure5-1.png",
        "caption": "Fig. 5 Development of viscous torques perpendicular to the journal spin axis",
        "texts": [
            "org/about-asme/terms-of-use In the region of low eccentricity ratios, the solutions to the linearized (first-order) equation are fairly good. Since the equation is linear, a linear sum of the two solutions (radial-load and torqueload) is itself a solution. The torque components, Mi and Mi, were determined by integrating the gas film pressure distribution. It was tacitly assumed that viscous torques acting about axes perpendicular to the spin axis were negligible in comparison to the pressure derived torques. Actually, differential viscous moments must arise (Fig. 5) because viscous shear stresses are inversely proportional to th e local gap width which is asymmetric when the bearing is mi saligned. The magnitude of these differential viscous stresses relative to the differential pressure stresses can be estimated aB follows: \u2022qoir Differential shear stress As c Hk c Differential pressure stress Ap epo 6 r Since Hk/6 is generally of the order of 1 while c/r<C 1, the viscous torques perpendicular to the spin axis may be neglected in comparison to the pressure derived torques"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002880_s40998-021-00421-0-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002880_s40998-021-00421-0-Figure3-1.png",
        "caption": "Fig. 3 Structural sketch of the machine set",
        "texts": [
            " The fault type and working state of bearings can be obtained from the output of the trained SVM. Although there are many types of motor vibration signals, they have certain similarities according to frequency and energy distribution. Over the past decade, the data of the Case Western Reserve University (CWRU) Bearing Data Center website has become a standard reference for many researchers to test various algorithms. Thus, the data of CWRU are used to verify the method proposed in this paper. The structural of test stand is shown in Fig. 3, which mainly contains a motor, a torque transducer and a dynamometer. The type of rolling bearing is 6205-2RS JEM SKF deep groove ball bearing. The sample frequency is 12 kHz. Single point fault is arranged on the bearing by spark erosion technique. The diameter of the fault point is 0.1778 mm, and the fault depth is 0.2794 mm. Four types of vibration signals, including normal state, inner ring fault, outer ring fault and rolling bearing fault are collected (Loparo xxxx). The VMD algorithm needs to determine the modal decomposition number K in advance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002428_slct.202001751-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002428_slct.202001751-Figure1-1.png",
        "caption": "Figure 1. Schematic illustration of geometrical twisting with the formation of TICT state.",
        "texts": [
            "[34] Although one emission is due to intramolecular charge transfer from N-end to cyano moiety (normal emission), the red shifted emission they proposed as anomalous one which preferentially originates from solvent induced reversal of excited state.[34] Grabowski et. al. were the first to assign this lower energy emission is due to a dynamic relaxation involving molecular twisting of dimethyl amino group coupled with an intramolecular electron transfer from the donor group (dimethyl amino) to acceptor group (cyano moiety).[35] Hence the term comes twisted intramolecular charge transfer (TICT) state. The schematic illustration of geometrical twisting with the formation of TICT state has been shown in Figure 1. Such TICT formation is facilitated in high polar solvent due to lowering of the formation energy barrier of TICT state.[28] Red-shifted separate emission band from TICT state along with ICT state in polar solvent was also observed by Rettig and co-workers for p-(N,N-dimethylamino) benzonitrile (DMABN).[28] Stabilization of TICT state may occurs by maintaining a large twist angel between donor and acceptor group through intermolecular H- bonding.[36] Enhancement of TICT emission band of different molecule on \u03b1-cyclodextrine Dr"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000171_0954406219878741-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000171_0954406219878741-Figure1-1.png",
        "caption": "Figure 1. Dual tilt-quadrotor photo.6",
        "texts": [
            " Also, the new concept has better performance in navigating indoor areas since translational forces are produced by tilting the individual thrust vectors. All these advantages make the tilt-quadrotor platforms very attractive for fault and failure recovery. Several research groups worked on designing and building the novel actuation of tilt-quadrotors. This resulted in a range of developed platforms driven by different requirements. Some of these platforms have single tilt axis, while few have dual axes tilting rotor. The platform shown in Figure 1 is developed by coauthor.6,7 It is a relatively large platform compared to the rest of the platforms in literature. This platform has a mass of 4 kg, with 2 kg payload capacity. The size and mass of this platform required twobladed carbon fibre propellers of size 15 5 in. The tilting capability is achieved by two servos: one for the push\u2013pull mechanism and another servo to rotate the arm. The software and hardware architecture are discussed by Ryll8 to implement the feedback linearisation-based control system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000818_ecce.2016.7855220-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000818_ecce.2016.7855220-Figure7-1.png",
        "caption": "Fig. 7 \u2013 Sankey diagram of loss amounts for laser welding and interlocking production chains.",
        "texts": [
            " In general, it is not possible to make a global statement or to give a recommendation on manufacturing processes for stator cores because the losses depend on many individual properties such as the target geometry or the position of joint connection. Therefore, it is imperative to have the possibility to make measurements with product-specific geometries like single teeth, made possible by the newly developed single-tooth tester. If there is a specific production chain which want to be investigated, the described measurement methods can also be used to point out each step\u2019s amount of losses. Two examples are given in Fig. 7 for the stamping-interlocking and the stamping-laser-welding process. This kind of consideration is useful to determine the most relevant production step which is responsible for a major amount of losses. If there is a demand of increasing the manufacturing quality with regard to electromagnetic properties of the stack, it is recommended to focus on the determined step. VI. REFERENCES [1] Bertotti, Giorgio, \"General properties of power losses in soft ferromagnetic materials,\" in Magnetics, IEEE Transactions on , vol"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000477_0954406216631781-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000477_0954406216631781-Figure2-1.png",
        "caption": "Figure 2. Physical model of the fluid: (a) the rotating domain and (b) the fixed domain.",
        "texts": [
            " The spindle cooling fan is a stamped product with six fan blades mounted on the spindle by four hexagonal head bolts. Outside the fan area, a shield was mounted for protection. There are four vents in the rear cover to let the cold air pass. The numerical model is simplified according to the real motorized spindle. The model can be divided into the fluid analysis model and the heat transfer model. Different computational domains have different parameter settings in fluid analysis. As shown in the Figure 2, the fluid analysis model is divided into two parts, one is the rotating domain around the fan area and another is the fixed domain, which is the rest of the fluid area. The rotating domain includes the cylinder that envelopes the fan and the interface between air and the fan. The rest of the fluid area is a fixed part for the fluid state calculations. It is important to simplify the model for the thermal analysis. In the simulation of the motorized spindle, the rotor and the stator can be simplified as thick cylinders with evenly distributed heat source"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002020_s12206-020-1030-6-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002020_s12206-020-1030-6-Figure11-1.png",
        "caption": "Fig. 11. Structure of the finger mechanism.",
        "texts": [
            " The peak rotation range occurs at 4l = 1.5 cm, where 4 3l l= . The same phenomenon also occurs in Fig. 7(d). This reveals that the grasping range could be optimized by designing an equal length of link AD and BC. In the presented study, the finger operation range could achieve a maximum of r\u03b8 = 236\u00b0 when 3 4l l= . Based on the motion characteristic of the finger mechanism, a prototype model is developed to investigate the effectiveness of the finger mechanism in bidirectional object grasping, as depicted in Fig. 11. The presented finger mechanism has a parameter set of 1 2 5l l l= = = 5.5 cm, 3 4l l= = 2 cm. The link AB and CD is designed as S type to avoid motion interference and obtain a maximum grasping range. In Fig. 11, the yellow transparent area denotes the overall grasping range of the finger mechanism. To achieve object grasping, two modular fingers have been fabricated and form a configurable gripper, as depicted in Fig. 12. Each finger mechanism is actuated through a closed loop tendon. And the two tendons are actuated by the wire wheel which mounted on the output shaft of the micro servo motor. In the present configuration, the gripper could grasp object from outside to inside as the wire wheel rotates clockwise"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003709_s0043-1648(96)07486-8-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003709_s0043-1648(96)07486-8-Figure4-1.png",
        "caption": "Fig. 4. Sketch of the contact model.",
        "texts": [
            " In the analysis of the normal compliance it is assumed that the roller\u2013washer contact can be divided into a number of disks. These disks are situated in planes parallel to the radial plane of the rollers. Furthermore, it is assumed that the shear effects between the disks can be neglected. Thismethodwas used byHarris [10], who studied the influence of misalignment on the fatigue life of cylindrical roller bearings. For a steel on steel line contact, the normal contact defor- mation is given by Palmgren [11]: 9/10Py5ds3.84=10 (2)8/10l The roller\u2013washer contact is subdivided into n disks, each with width w (see Fig. 4), then lsnw (3) where l is the roller effective contact length i.e. the part of the roller that is in contact with the washer. The load per unit length can be expressed as P ps (4) l Substituting Eq. (4) into Eq. (2) gives the following expression: 10/9d ps (5)y5 1/91.24=10 (l) Since a specific load is required in the analysis, an estimated contact deformation, d, is used as starting value. The load per unit length, p, as well as the effective contact length, l, is calculated several times in an iterativemanner, until the specified load is achieved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000844_1464419318819332-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000844_1464419318819332-Figure4-1.png",
        "caption": "Figure 4. Radial and thrust loading effect on roller by Harris and Kotzalas.40",
        "texts": [
            " Deflection of IR 2pl and OR 1pl interaction with roller is denoted by 1pl \u00bc 1p b l 1 2 p \u00fe q 2 l 1 2 q\u00fe 1\u00f0 \u00de b dm \u00fe d \"2p Z 2 Bd \u00f09\u00de Bdo \u00bc r r cos J 2r sin o J mod di, 2 \u00f0 \u00de \"\u00f0 \u00de , 04mod di, 2 \u00f0 \u00de \"5 d1 r r cos J 2r , d14mod di, 2 \u00f0 \u00de \"5 d d1 r r cos J 2r sin o J mod di, 2 \u00f0 \u00de \"\u00f0 \u00de , d d14mod di, 2 \u00f0 \u00de \"5 d 8>< >: \u00f07\u00de Bdi \u00bc r r cos J 2r : sin i J mod di, 2 \u00f0 \u00de\u00f0 \u00de , 04mod di, 2 \u00f0 \u00de5 d1 r r cos J 2r , d14mod di, 2 \u00f0 \u00de5 d d1 r r cos J 2r : sin i J mod di, 2 \u00f0 \u00de\u00f0 \u00de , d d14mod di, 2 \u00f0 \u00de5 d 8>< >: \u00f06\u00de 2pl \u00bc 2p \u00fe b l 1 2 p \u00fe l 1 2 q\u00fe 1\u00f0 \u00de b 2 dm \u00fe d \"2p Z 2 Bd \u00f010\u00de where Bd is an amplitude either due to RD or due to combined IRD\u2013ORD\u2013RD. The effect of radial and thrust load on the roller is depicted in Figure 4. State of balancing for deflection. Deflection in axial and radial directions, radial gap between roller and races, and highest total deformation for IR and OR is used for the equilibrium equation of deflection a g d h i \u00fe r cos p Z 2 \" 1p b l 1 2 p \u00fe q 2 l 1 2 q\u00fe 1\u00f0 \u00de b dm \u00fe d \"2p Z 2 Bd # max \u00fe 2p \u00fe b l 1 2 p \u00fe l 1 2 q\u00fe 1\u00f0 \u00de b 2 dm \u00fe d \"2p \" Z 2 Bd max \u00bc 0 \u00f011\u00de where p \u00bc angle of pth roller\u00bc 2 p 1\u00f0 \u00de N \u00fe ct State of balancing for radial load. For radial load equilibrium condition 0:62 10 5Qr b0:89 Xp\u00bcN 2\u00fe1 p\u00bc1 vp: cos p q0:112p Xl\u00bcq l\u00bc1 2p \u00fe b l 1 2 p \u00fe l 1 2 q\u00fe 1\u00f0 \u00de b 2 dm \u00fe d \"2p Z 2 Bd #1:11 \u00bc 0 \u00f012\u00de where vp \u00bc load zone defined by Harris and Kotzalas"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000748_978-981-10-2875-5_114-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000748_978-981-10-2875-5_114-Figure5-1.png",
        "caption": "Fig. 5 Linear-play operator",
        "texts": [
            " Assuming that the volume of PMA is constant during charging and discharging, then P _V 0 and the internal pressure of PMA can be obtained by taking integration of Eq. (7). In this section, the Prandtl-Ishlinskii (P-I) model [17] will be used to derive the hysteresis of the PMA. It has two advantages: first, it is simpler compared with other models because it only consist of linear play operators; second, the inverse PI model can be obtained analytically, which is easier for realization of hysteresis compensation. The elementary operator of the PI model is linear play operator, which can be mathematically illustrated by Fig. 5. Its ith linear play operator can be expressed as yi\u00f0k\u00de \u00bc max x\u00f0k\u00de ri;min x\u00f0k\u00de\u00fe ri; yi\u00f0k 1\u00def gf g \u00f08\u00de while the initial condition is yi\u00f00\u00de \u00bc max x\u00f00\u00de ri;min x\u00f00\u00de\u00fe ri; yi0f gf g \u00f09\u00de where Hr denotes the linear play operator; x = [x1, \u2026, xn] T is the weighting vector; r = [r1, \u2026, rn] T is the threshold vector; x and y are the input and output of the operator, respectively; y0 is the initial state; k is the sampling number of the operator; and n is the number of the linear play operator. To determine the parameters of PI model, the threshold vector r is firstly determined by the following equation ri \u00bc i n\u00fe 1 maxf x\u00f0t\u00dej jg i \u00bc 1; "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001782_j.vacuum.2020.109557-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001782_j.vacuum.2020.109557-Figure2-1.png",
        "caption": "Fig. 2. Schematic illustration of (a and b) SLM system and (c) hot isostatic pressing system (T: temperature; P: pressure; CR: cooling rate).",
        "texts": [
            " The hot isostatic pressing (HIPing) was performed on an AIP14-30H facility (AIP, United States), equipped with a maximum temperature of 2000 \ufffdC and maximum gas pressure of 207 MPa, for the densification experiment. High pure argon gas was used as protection and pressure medium to avoid the undesired oxidation. The temperature, pressure, holding time and heating/cooling rate were set as 950 \ufffdC, 150 MPa, 120 min, and 10 \ufffdC/min, respectively. After the HIPing process, the relative density of SLMed sample remains high value without clear change. The manufacturing process in this work is schematically illustrated in Fig. 2. The phase was characterized by X-ray diffraction (XRD, PANalytical, Netherlands) with a scanning speed about of 10\ufffd/min, which equipped with Co radiation operated at 40 kV and 40 mA. Metallographic samples were prepared by mechanical and OP-S polishing. A chemical solution of N. Kang et al. Vacuum 179 (2020) 109557 1 ml HNO3 and 24 ml CH5OH was used as the etching agent. The oxygen content is determined using a LECO Corporation ONH836 instrument (Germany). Then, the microstructure was characterized by scanning electron microscopy (SEM, Gemini 500, Zeiss Germany)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002628_ac60119a012-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002628_ac60119a012-Figure1-1.png",
        "caption": "Figure 1. Effect of nitrous acid on absorption spectrum Figure 2. Effect of nitrous acid on absorption spectrum af p-phenylenediamine at pH 1.4 of p-amino-N-methylacetanilide at pH 1.1",
        "texts": [],
        "surrounding_texts": [
            "Apparatus. Absorption spectral data were obtained on three different instruments. A Cary Model 10-1 1M recording spectrophotometer mas used in all the preliminary work nhere the entire ultraviolet spectrum mas wanted. A11 calibration curves, interference effects, and stability and precision studies were obtained using the Beckman Model B and Model DU spectrophotometers. The sensitivity control on the Beckman Model B was kept a t 2 for this work. hlatched 1-cm. quartz absorption cells were used with all the instruments. Glass-stoppered volumetric flasks of 50-ml. capacity were used in the preparation of all of the solutions run on the instrument. All pH measurements were made on a Beckman Model H-2 glass electrode pH meter. Reagents. The sodium nitrite, potassium permanganate, sodium thiosulfate, and potassium iodide used in the preparation and standardization of the stock nitrite solution were Baker\u2019s analyzed reagents. All the p-phenylenediamine derivatives investigated were the best grade obtainable of Eastman organic, chemicals. The amine hydrochlorides were used whenever available because of the greater stability of the acid salts compared to free amines. The Eastman catalog number of the various reagents is included in the data of Table I. Stock solutions of the various amines investigated were prepared in 0.005- or 0.01X concentration by dissolving an accurately weighed portion of the reagent in distilled v-ater acidified with 1 to 4 hydrochloric acid, followed by dilution to the desired volume. The final acid concentration of the stock solutions used in the studies of the effect of pH and reagent conStock Solutions. V O L U M E 28 , NO. 11, N O V E M B E R 1 9 5 6 1675 - 0.0008.M solution of reagent - - - 0.0008-W solution of reagent 7 to 1: minutes after addition of 0.178 mg. of nitrite ion per 50 ml. - 0.0002A4 solution of reagent 0.0002.~ splution of reagent 10 to 15 minutes after addition of 0.059 mg. of nitrite ion per 50 ml. _ _ _ centration was approximately 1 to 100 or 1 to 50 hydrochloric acid ( 5 or 10 ml. of 1 to 4 hydrochloric acid per 100-ml. volume). Stock solutions of the amines prepared for the calibration work were usually relatively acidic (1 : 12 to 1 : 4 hydrochloric acid) solutions. The acid content of each stock solution was so adjusted that the aliquot most frequently takeii would contain sufficient acid to give the desired pH (1.0 to 1.5). This eliminated the need for a separate addition of acid in preparing the test solutions to be run on the instruments. Stability studies on the reagents themselves were carried out on these acidified stock solutions, as they were generally more stable than those of the same concentration of reagent prepared in more dilute acid. The standard nitrite solutions used throughout this work \\yere approximately 0.00025M (0.0115 mg. of nitrite ion per ml.), and were prepared by successive dilution of 0.05- and 0.01X solutions of sodium nitrite. The 0.0531 nitrite solution was standardized titrimetrically, by adding an excess of potassium permanganate to oxidize the nitrite, followed by a thiosulfate titration of the iodine liberated by the action of the excess permanganate on potassium iodide. Details of this standardization procedure are givrn bv Kolthoff and Sandell ( 3 ) . The exact concentration of the 0.00025J1 nitrite solution was calculated from the standard value of the 0.0551 solution, using the necessary dilution factors. SURVEY OF PHENYLEKEDIAMINES Effect of Nitrous Acid on Absorption Spectra. Figures 1 to 3 illustrate the type of absorption changes which takes place upon the addition of nitrite ion to phenylenediamine derivatives in arid solution. rlll of the compounds investigated give very similar changes. The principal absorption maxima of all the diazoilium salts occur in the 330- to 380-nip region, whereas the reagents themselves shon- no significant absorption. Table I summarizes the ultraviolet spectral data for the various reagents. The data apply to systems having pH's of 1.0 to 1.4, unless otherwise stated. The sensitivities of the various reagents toim,rd nitrite i o n are expressed as molar absorbancy indices (a11 values) These values xere calculated according to the equation, a.11 = A.4.(2300)/c, where A A , is the absorbancy change of the system a t a specific wave length upon the addition of c nig. of nitrite ion per 50 ml. and 2300 is a factor converting the nitrite concentration to moles per liter. The ultraviolet cutoffs of the reagents themselves are also listed in Table I. The nave length a t which the reagent starts to absorb significantly (A48 = 0.01) is given, together with the reagent concentration aiid pH of the solution to rrhich the cutoff applieq. The diazonium salts formed from the p-phenylenediamines are light-sensitive. If no precautions are taken to shield the solutions from light after mixing, the diazonium salts decompose a few minutes after they are formed, as is el-idenced by the rapid decrease in the intensity of the Stability of the Systems. - 0.OOl.M solution of reagent _ _ _ 0.001.M solution of reagent 3 to 23 minutes after addition of 0.059 mg. of nitrite ion per 50 ml. absorption maxima in the 330- to 380-mp region. The absorbancy obtained depends upon how long the solutions m r e exposed to the light before they rrere placed in the spectrophotometer. If the solution is transferred to the absorption cell and placed in the cell compartment immediately after mixing, an extremely stable system results. Under these circumstances, the absorbancy quickly rises to a maximum and remains constant if the solution is not exposed to the light. All the work reported herein pertains to solutions placed in the instrument immediately after mixing. The time intervals specified in Table I pertain to the time after mixing during which the systems n-ere found to give a constant absorbancy reading, if kept in the darkness of the spectrophotometer cell compartment. S o formal stability studies of the various systems n-ere undertaken for periods longer than those specified. However, i t was apparent during the course of the investigation that the majority of the systems are stable for hours if light is excluded. The effect of pH and reagent concentration of the system was determined in some instances. Varying the pH in the range 1.0 to 2.4 usually changed the spectrum of the reagent considerably, but had little or no effect on the intensity of the absorption maximum of the diazonium salt in the 330- to 380-mp region. I n general, it was observed that both the formation and light-catalyzed decomposition of the diazonium salt proceed a t slower rates as the pH is raised. Perchloric and sulfuric acid were substituted for hyEffect of pH and Reagent Concentration, 1676 A N A L Y T I C A L C H E M I S T R Y The Model DU spectrophotometer was calibrated for the determination of nitrite ion using several of the reagents. Details of the calibration procedure will be found in the following section. The phenylenediamine, chlorophenylenediamine, 2,5-toluenediamine, and N,N-diethylphenylenediamine systems all followed Beer's law a t the wave lengths given in Table I up to the calibration limit of 0.06 mg. of nitrite ion per 50 ml. Slit widths of 0.08 to 0.11 mm. m r e used in this work on the Model DU. The Beckman Model B spectrophotometer was also calibrated using a few of the reagents. The aminoacetanilide system follows Beer's law over the range 0.00 to 0.08 mg. of nitrite ion per 50 ml. when this instrument is used. The N,N-diethylphenylenediamine system s h o w a slight deviation from the usual straightline plot above 0.04 mg. of nitrite ion per 50 ml., in contrast to its behavior using the Model DU. Calibration Curves. drochloric acid in some of the tests involving phenylenediamine and 2,8toluenediamine. The use of these acids resulted in s y s terns which gave less intense absorption maxima, and xhich reached these maxima more slowly than the corresponding systems in hydrochloric acid. The concentration of several of the reagents was varied from 0.0002- to 0.001- or 0.002M, with negligible effects upon the intensity of the absorption maxima of the diazonium salts. 250 275 300 325 350 375 W A V E L E N G T H ( m p l F igure 4. Effect of n i t r o u s acid on absorpt ion spec t rum of chloro-p-phenylenediamine at pH 1.2 - 0.001-W solution of reagent _ - - 0.001M solution of reagent 3 to 11 minutes after addition of 0.038 nig. of nitrite ion per 50 ml. Stability of the Reagents. Aromatic amines in general are sensitive to air and light, and should be protected therefrom as much as possible. The amine hydrochlorides were chosen for this work, when available, so as to minimize the decomposition n-hich often occurs in amino compounds upon prolonged standing. Stability studies were carried out on hydrochloric acid solutions of several of the reagents. A 0.0075M solution of phenylenediamine in 1 to 4 acid and a 0.05M solution in 1 to 50 acid were found to be stable for 48 hours. The less acidic solution showed a more rapid rate of decomposition after this period. A 0.01M stock solution of S,N-diethylphenylenediamine in 1 to 12 acid was found to be stable for 7 days after its preparation. I t gave reproducible absorbancy readings upon the addition of nitrite ion during this period, despite the development of a slight pink hue in the solution after 3 or 4 days. A 0.001M solution of aminoacetanilide in 1 to 50 acid gives reproducible absorbancy readings for 5 days after its preparation. Stock solutions 0.OlM in 2,5-toluenediamine, in both 1 to 100 and 1 to 4 hydrochloric acid, showed signs of decomposition 2 to 3 days after their preparation. The similarity in absorption characteristics among the various systems investigated is apparent from Table I. Compounds having N-substituted amino groups give the same type of absorption maximum, in the same region and of comparable intensity to those having two free amino groups. This would indicate that only one of the free amino groups is involved in the reaction with nitrite ion, and that diazotization and not tetrazotization is the predominant reaction. To confirm this belief, mole ratio studies were carried out on the phenylenediamine, chlorophenylenediamine, and 2,5,-toluenediamine systems. The data in Figure 5 pertaining to chlorophenylenediamine are indicative of the type of plot obtained in these studies, and the range of concentrations employed. Values of 0.92 and 0.87 mole of nitrite per mole of reagent were obtained for the phenylenediamine system. 2,%Toluenediamine gave 0.90 and 0.82 mole of nitrite per mole of reagent. Results of the mole ratio studies using chlorophenylenediamine are discussed in the following section. Mole Ratio Studies."
        ]
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure4-1.png",
        "caption": "Fig. 4a. Continuous Core Ring",
        "texts": [
            " But this effect was observed in a few unusual cases where the coupling was underloaded due to its being large for the duty, or where the horse-power required by the fan was less than had been estimated. I t was soon realized that the liquid filling the voids formed between the two semicircular core guide rings would not be effective in the transmission of power, and that the slope of the quantity-slip curve would be improved by closing these voids. The usual practice is to fit the annular filling plates shown in Fig. 4a, Plate 1. Typical quantity-slip curves for a scoop tube coupling driving a fan load are shown in Fig. 2. Curve A relates to a standard \u201cVulcan\u201d circuit with core ring and filling plates working at a specific load normal for such a fan drive, and curve B represents an underloaded condition which might reasonably be encountered in practice. Although the two curves are not uniform in shape no flat spot is present and no difficulty is experienced with hand or automatic control in adjusting the quantity and speed to the precise value required",
            " Note that the quantity of liquid may actually be varied over a range of 12 per cent under these conditions without changing the speed. Interrupted Core Ring. The correct solution of the \u201cflat spot\u201d problem appears obvious now that it has been found. It merely involves 0 25 SO 75 FILLING OF WORKING CIRCUIT-PER C E N Fig. 2. Quantity-Slip Curves for Continuous Core Ring Coupling (Fan Drive) Fig. 3. Quantity-Slip Curves for Interrupted Core Ring Coupling (Fan Drive) cutting ports through, or interrupting sections of, the core guide ring at regular intervals, as shown in Fig. 4, Plate 1. The impeller and runner in the upper photograph have a standard \u201cVulcan\u2019) circuit with continuous core ring, and those below have the interrupted core ring. Three progressive degrees of filling of a Vulcan circuit which is oversized for the power to be transmitted are illustrated by Fig. 5, and the extent to which the core guide ring causes the \u201cflat spot\u2019) by interfering 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 87 with the progressive building up of the vortex ring is clear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001752_j.mechmachtheory.2020.103992-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001752_j.mechmachtheory.2020.103992-Figure12-1.png",
        "caption": "Fig. 12. Finite element model of the pinion and the gear.",
        "texts": [
            " Similarly, the cross-section curves of the convex side can be set to be the same as those of the concave side. In the convex side, K (12 u ) is set as 0.236 mm \u22121 , and R (1) = R (2) = 8.4 mm. Finally, from Eqs. (8) \u2013(12) , the tooth surface of the pinion and the gear can be exported, and the solid models of the pinion and the gear can be built as shown in Fig. 11 . 5.5. Stress analysis Based on finite element analysis, a loaded contact analysis can be completed to evaluate the stress distributions of the pinion and gear in Fig. 11 . The finite element model is as shown in Fig. 12 , which is meshed by the first-order hexahedral element. For the elements on the tooth surface, the max size of element length is set as 0.2 mm, and for other elements, the max size is set as 2 mm. The Young\u2019s modulus is 2.1 \u00d7 10 11 Pa and Poisson\u2019s ratio is 0.267. The torque of 540 Nm applied on the gear and the pinion is fixed as shown in Fig. 13 . The results of finite element analysis are shown in Fig. 14 . It can be seen that the shape and orientation of the contact ellipse of the finite element analysis are similar to those of the analytical analysis in Section 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000763_s10443-016-9565-5-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000763_s10443-016-9565-5-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of warp-reinforced 2.5D woven composite and definition of RVC",
        "texts": [
            " Then, the predicted results are shown and compared with experiment in Section 6. Finally, some valuable conclusions are summarized in section 7. The simulation precision of the mechanical behavior of textile composites is highly dependent upon their actual microstructure. The analytical geometry model proposed in this paper takes into account the inherent geometric complexity, including the cross-section and curvatures of yarns. The weave style considered in this paper is a type of 2.5D woven reinforcement: warpreinforced 2.5D as shown in Fig. 1. The warp and weft are oriented along the x-direction and y-direction, respectively. The binder warp interlocks with the weft layers through the thickness (z-direction). Also, the warp and binder warp arranges alternately with a certain ratio of 1:1 along the x-direction. Therefore, in this structure, warp remains straight to the maximal degree so as to make the composite achieve higher in-plane mechanical properties. The architecture of the undeformed composite is periodical in warp, weft and thickness directions. The Representative Volume Cell (RVC) is marked with black box in Fig. 1, as well as its basic dimensions: A, B and C The preform structural parameters which are also known as manufacturer specified parameters are the number of layers (n), the warp density (Pwarp), the weft density (Pweft), the blinder warp density (Pblinder), the warp fineness (Texwarp), the weft fineness (Texweft) and the blinder warp fineness (Texblinder). These detailed parameters are summarized in Table 1. The material properties are summarized in Table 2. On the basis of the microscopic image analysis (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001475_1077546320903195-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001475_1077546320903195-Figure2-1.png",
        "caption": "Figure 2. Viscoelastic beam of length (L), cross-sectional area A \u00bc WH.",
        "texts": [
            " This kind of study, indeed, is useful for a first understanding of the physical parameters enclosed in the problem. Then, two relaxation times are taken into account, and their influence on the dynamics of the beam is deeply evaluated and described. Finally, some considerations are pointed out regarding the vibrational response of the beam in case of real viscoelastic materials. In this section, the analytical dynamic response of a viscoelastic beam with rectangular cross section is derived. Let L, W , andH , respectively, be the length, width, and thickness of the beam (Figure 2), and let us assume that L W , L H . Assuming also that the displacement along the z-axis ju\u00f0x; t\u00dej L, the Bernoulli theory of transversal vibrations can be applied, and therefore, it is possible to neglect the influence of shear stress in the beam. It is worth noticing that this hypothesis does not limit the validity of the analysis because attention is paid to the first resonant peaks, which are not affected by shear deformations. Hence, the general equation of motion is (Inman, 1996) Jxz Z t \u221e E\u00f0t \u03c4\u00de \u2202 4u\u00f0x; \u03c4\u00de \u2202x4 d\u03c4 \u00fe \u03bc \u22022u\u00f0x; t\u00de \u2202t2 \u00bc f \u00f0x; t\u00de (3) where \u03bc \u00bc \u03c1A, where \u03c1 is the bulk density of the material that the cantilever is made of and A is the area of the cross section of the beam, i",
            " With regard to the transverse motions of a narrow, homogenous beam with a bending stiffness E0Jxz and density \u03c1, the value of the natural frequencies can be calculated using a simple formula which is always valid, regardless of the boundary conditions (Thomson and Dahlelh, 1997) \u03c9n \u00bc cn L 2 ffiffiffiffiffiffiffiffiffiffi E0Jxz \u03c1A s (26) where coefficient cn depends on the specific boundary conditions. The first natural frequency, in particular, can be written as \u03c91 \u00bc \u03b12\u03b41 (27) being \u03b41 \u00bc c21 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi E0A=\u00f0\u03c1Jxz\u00de p and \u03b1 \u00bc Rg=L the dimensionless beam length, with Rg \u00bc ffiffiffiffiffiffiffiffiffiffi Jxz=A p the radius of gyration. For the rectangular beam cross section under investigation (Figure 2), one has \u03b1 \u00bc H=\u00f0L ffiffiffiffiffi 12 p \u00de and \u03b41 \u00bc \u00f0c21=H\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffi 12E0=\u03c1 p . The nondimensional eigenvalue is now defined as s \u00bc s=\u03b41 (28) and in particular one has, for the nth mode, \u03c92 n \u00bc E0\u03b2 4 nJxz=\u03bc \u00bc rnE0 and \u03b4n \u00bc c2n ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi E0A=\u00f0\u03c1Jxz\u00de p . By substituting equation (28) into (25), the following nondimensional characteristic equation is obtained s3 \u00fe s2 1 \u03b81 \u00fe \u00f01\u00fe \u03b31\u00de\u03b14\u03942 ns\u00fe 1 \u03b81 \u03b14\u03942 n \u00bc 0 (29) where \u0394n \u00bc \u03b4n=\u03b41, and having defined the dimensional groups \u03b81 \u00bc \u03b41\u03c41 (30) \u03b31 \u00bc E1=E0 (31) Equation (29) can be then rewritten as s3 \u00fe X2 j\u00bc0 ajjs j \u00bc 0 (32) where a0 \u00bc \u00f01=\u03b81\u00de\u03b14\u03942 n,a1 \u00bc \u00f01\u00fe \u03b31\u00de\u03b14\u03942 n, and a2 \u00bc 1=\u03b81",
            " Moreover, it is possible to define, for the quartic equation (39), the discriminant D2\u00f0n\u00de (Lazard, 1988; Rees, 1922) D2\u00f0n\u00de \u00bc 256a30 192a3a1a 2 0 128a22a 2 0 \u00fe 144a2a 2 1a0 27a41 \u00fe 144a23a2a 2 0 6a23a 2 1a0 80a3a 2 2a1a0 \u00fe 18a3a2a 3 1 \u00fe 16a42a0 4a32a 2 1 27a43a 2 0 \u00fe 18a33a2a1a0 4a33a 3 1 4a23a 3 2a0 \u00fe a23a 2 2a 2 1 (41) which can be used to deduce important properties of the roots of equation (39). The beam cross-sectional acceleration A\u00f0x; s\u00de is in this case A\u00f0x; s\u00de \u00bc F0 X\u00fe\u221e n\u00bc1 s2\u00f01\u00fe \u03b81s\u00de\u00f01\u00fe \u03b82s\u00de\u03c6n\u00f0x\u00de\u03c6n xf \u03bc\u03b81\u03b82 s4 \u00feP3 j\u00bc0ajs j (42) In this section, the main results of the presented analysis are discussed. The flexural vibrations of a viscoelastic beam with rectangular cross section and thickness H \u00bc 1 cm, which oscillates in the xz-plane (Figure 2), are studied. The only geometrical parameter that is considered varying in calculations is the beam length L. In particular, the ratio \u03b1 \u00bc Rg=L is changed maintaining Rg \u00bc H= ffiffiffiffiffi 12 p constant. At first, quantitative investigation of a specific viscoelastic material is not carried out, but a qualitative study of a generic viscoelastic beam behavior, which can be considered at different lengths L (e.g. in experimental testing campaigns, to cover wide frequency ranges) and at different working temperatures (i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003630_047134608x.w1111-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003630_047134608x.w1111-Figure4-1.png",
        "caption": "Fig. 4. Simplified representation of three one-dimensional turbulence fields.",
        "texts": [
            " Turbulence can be defined as a particular type of nonquiescent atmosphere in which rapid fluctuations of wind velocity and direction occur. An introductory treatment of this phenomenon and the manner in which it is modeled is given by Hess (4). For the purposes of this exposition, turbulence will be modeled in its simplest form. It will be assumed that the aircraft can encounter three one-dimensional turbulence velocity fields, wherein the air-mass velocity has three components ug, wg, and wg parallel to the aircraft\u2019s x, y, and z body axes in the equilibrium condition. Figure 4 indicates these three velocity fields. The amplitude of each component will vary with time, but is assumed invariant over the length of the aircraft and is equal to the amplitude experienced at the aircraft\u2019s center of gravity. Equations (16) and (17) can be modified to accommodate this simple turbulence model by realizing that: (1) The aircraft will see a relative wind with perturbation components u \u2212 ug, v \u2212 vg, and w \u2212 wg. (2) The effects of this simplified turbulence representation on the dynamics of the aircraft can be approximated by modifying the aerodynamic (as opposed to the inertial) terms in Eqs",
            " With a control system like that just outlined in operation, an aircraft can be forced to follow a desired trajectory in three-dimensional space at some desired speed. In terms of mission effectiveness or economy some trajectories can be considered more desirable than others. For example, consider the fighter aircraft of Fig. 1 in a combat situation. To avoid an adversary the pilot must change heading (Euler angle ) by 180\u25e6 as quickly as possible. The question is: What trajectory and speed profile will allow this heading change to occur in minimum time? Next consider a large passenger aircraft such as that shown in Fig. 4. To minimize ticket costs it is necessary to minimize fuel consumption. The question now is: What trajectory and speed profile from departure point A to destination point B will result in minimum fuel consumption? The answers to these questions involves the discipline of trajectory optimization. The mathematical tools involved are those of the calculus of variations. The equations describing the aircraft motion are typically simpler than those just derived. The aircraft is represented as a point mass, upon which act the gravitational, aerodynamic, and propulsive forces",
            " An instructive starting point for a discussion of aircraft control design is the so-called classical approach using fundamental analysis and synthesis tools such as the Laplace transform and Bode diagram in a process referred to as loop shaping. The Bode diagram is a plot of the following functions: where gij(s) denotes any of the transfer functions of Eq. (25). The left-hand side of the first of Eqs. (28) defines the magnitude in decibels (dB). An Example. A simple classical design example will now be presented. The vehicle chosen is a jetpowered transport similar to that shown in Fig. 4. A single flight condition consisting of cruising flight at Mach 0.84 and altitude 33,000 ft has been chosen for study. Only longitudinal motion will be examined, and the control inputs will be the elevator angle \u03b4e and engine thrust \u03b4T. The synthesis tools to be employed can be found in any undergraduate text on the subject, for example, Nise (9). Table 1 lists the stability derivatives for this flight condition. Referring to Eqs. (19), any stability derivatives appearing in the equations but absent from Table 1 are assumed negligible"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000529_j.apm.2016.03.048-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000529_j.apm.2016.03.048-Figure2-1.png",
        "caption": "Fig. 2. Geometry of a curvilinear thrust bearing.",
        "texts": [
            " If the values of \u03bc do not vary, the non-linearity of the flow curve increases with the value of k , which stands for the coefficient of pseudo-plasticity. In pseudo-plastic fluids k = 0 and in Newtonian fluids k = 0 . Therefore, in Newtonian fluids, the initial viscosity becomes the viscosity given by Newton\u2019s law. The three-dimensional notation of Eq. (2.2) may be expressed as [10] : \u03bcA 1 = ( 1 + k 2 ) where = [ 1 2 tr ( 2 )] 1 2 , (2.3) is the magnitude of the second-order shear stress tensor , but A 1 is the first Rivlin\u2013Ericksen kinematic tensor. Let us consider a thrust bearing with a curvilinear profile of the working surfaces shown in Fig. 2 . The lower fixed surface is described by the function R ( x ) which denotes the radius of this surface. The nominal bearing clearance thickness is given by the function h ( x, t ). The expression for the film thickness is considered to be made up of two parts: H = h ( x, t ) + h s ( x, \u03d1, \u03be ) , (2.4) where h ( x, t ) represents the nominal smooth part of the film geometry, while h s = \u03b4r + \u03b4s denotes the random part resulting from the surface roughness asperities measured from the nominal level, \u03be describes a random variable which characterizes the definite roughness arrangement. An intrinsic curvilinear orthogonal coordinate system x , \u03d1, y linked with the upper surface of a porous layer is also presented in Fig. 2 . Taking into account the considerations of the works (Walicka [10] ; Walicki [11] ) the equation of continuity and the equations of motion of a R\u2013R\u2013S fluid for axial symmetry can be presented in the form: 1 R \u2202 ( R \u03c5x ) \u2202x + \u2202 \u03c5y \u2202y = 0 , (2.5) Please cite this article as: A. Walicka et al., Curvilinear squeeze film bearing with rough surfaces lubricated by a Rabinowitsch\u2013Rotem\u2013Shinnar fluidAU: Please validate the article title., Applied Mathematical Modelling (2016), http://dx.doi.org/10.1016/j"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure4-1.png",
        "caption": "Fig. 4. Configurations of Mechanism I in circuit 1: (a) Configuration A: h1 = \u2212p/2, (b) configuration B: h1 = 0, and (c) configuration C: h1 = p/2.",
        "texts": [
            " 4 and 5 show Mechanism I at configurations A, B and C in circuit 1 and configurations D, E and F in circuit 2 respectively. It is observed that at two configurations in each circuit (see configurations A and C in circuit 1 and configurations D and F in circuit 2), the axes of R joints 1, 3, 4 and 6 of Mechanism I are coplanar and the axes of R joints 2 and 5 are perpendicular to the plane defined by the axes of R joints 1, 3, 4 and 6. The above results have been verified using several mechanism models built using 3D printing. Fig. 6 shows the CAD model and 3D-printed prototype of 6R Mechanism I (Fig. 4c). It is noted that joints 1 and 6 in this prototype are prevented from full-cycle rotation due to interference between links 2 and 4 as well as links 1 and 5. Let K1, K2 and K3 denote the intersections of joint axes of joints 1 and 6, joints 2 and 5, and joints 3 and 4. P1, P2 and P3 represent the plane defined by the axes of joints 1 and 6, joints 2 and 5, and joints 3 and 4 respectively (Fig. 7). From Ref. [36], we obtain that planes P1,P2 and P3 and plane K1K2K3 have a common point, K, at any configuration of the 6R mechanism during motion. It is noted that at configurations A, C (Fig. 4), D and F (Fig. 5) of Mechanism I, point K2 is at infinity. A 6R mechanism that has three pairs of R joints with intersecting joint axes has been proposed using a geometric construction approach. Kinematic analysis of the mechanism has been presented. The analysis has shown that the 6R usually has two solutions to the kinematic analysis for a given input. In two configurations in each circuit of the 6R mechanism, the axes of four R joints are coplanar, and the axes of the other two R joints are perpendicular to the plane defined by the above four R joints"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001207_rpj-07-2018-0171-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001207_rpj-07-2018-0171-Figure8-1.png",
        "caption": "Figure 8 (a) Testing device \u2013 scheme with pickup points and forces, (b) photogrammetry testing (both in braking load case)",
        "texts": [
            " During post-processing, the component was cut out from the platform, the support structures were manually removed and the component was sandblasted (Figure 7). Finally, the mounting surfaces were milled using the computer numerical control machine at the external workplace (S&K Tools s.r.o., Brno,CzechRepublic). The dimensional accuracy of the axle carrier after additive manufacturing was checked by 3D scanning using the ATOS Triple Scan (GOMGmbH, Braunschweig, Germany). A special testing device was developed for the verification of real component deformation. It is described using Figure 8(a) in the configuration of the breaking load case. The main part of the testing device is a welded frame that allows the fixation of the axle carrier [blue in Figure 8(a)] at the suspension pickup points (A, B, C points in Figure 1). The loading is applied through the welded beam substituting real formula wheel [violet part in Figure 8(a)]. The beam is connected to the axle carrier similarly as the wheel using the bearings and brake Topologically optimized axle carrier Ond rej Vaverka, Daniel Koutny and David Palousek Rapid Prototyping Journal Volume 25 \u00b7 Number 9 \u00b7 2019 \u00b7 1545\u20131551 calliper. At the estimated wheel diameter, the loading forces (FG and FB) represented by threaded rods are applied. Loading forces are applied by tightening of the nuts in the direction of the arrows, as shown in Figure 8(a), whereas the loading forces are directlymonitored by strain gauges. The deformations during loading were captured with the photogrammetry systemTRITOP (GOMGmbH, Braunschweig, Germany). The axle carrier and the frame were covered with photogrammetry targets during the loading process. Two types of targets were used \u2013 3mm (frame) and 1.5mm (axle carrier). The whole scene [Figure 8(b)] was recorded using the standard capturing method, and then the focus of the camera was changed, and 18 detailed photos of the axle carrier were made for the bestpossible accuracy (three images in each of three height levels from both sides). Deformations were captured for 0, 40, 80, 100 and 120 per cent of loading forces. Deformations obtained from photogrammetry measurement were then compared with those fromFEManalysis. According to FEM analysis, maximal deformation of the axle carrier was 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000325_3387304.3387327-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000325_3387304.3387327-Figure4-1.png",
        "caption": "Figure 4. F450 Multicopter Model in Gazebo",
        "texts": [
            " In each episode, a set of action, a state, a reward is saved for an RLNN. Using these sets, MGAT learns how to get a better reward than the latest learning and updates a neural network in every 10th episode. N RESULTS Before doing the simulation, it is needed to apply the vehicle we use in it. Through the motor thrust test and the multicopters' specifications, we implemented the following. We performed it in GAZEBO World to adapt MGAT, and the following is the result. This experiment only confirmed the tendency without prior learning (Fig. 4-6). Through this, we were able to find out the possibility of the autotuning algorithm and its applicability to the actual vehicle. The controller has shown that the reactivity about the angle has reached the ideal target. If the experiment goes further, we will get more ideal gains value. We did not put the thrust control into the internal MGAT algorithm, and it would be necessary to have a very large space for the actual flight in case it would fall. The result of the operation of this algorithm in TX2 will be shown in the following results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002032_j.engfailanal.2020.105082-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002032_j.engfailanal.2020.105082-Figure2-1.png",
        "caption": "Fig. 2. Finite element contact model of herringbone gears.",
        "texts": [
            " Engineering Failure Analysis xxx (xxxx) xxx \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 \u03c3x = \u2212 2z \u03c0 \u222b a \u2212 a P(x)(x1 \u2212 x)2 [(x1 \u2212 x)2 + z2 1] 2 dx \u03c3z = \u2212 2z3 \u03c0 \u222b a \u2212 a P(x) [(x1 \u2212 x)2 + z2 1] 2 dx \u03c4xy = \u2212 2z2 \u03c0 \u222b a \u2212 a P(x)(x1 \u2212 x) [(x1 \u2212 x)2 + z2 1] 2 dx (3) Eq. (2) is substituted into Eq. (3). The maximum von Mises stress depth is 0.7a and the maximum shear stress depth is 0.5a. Finite element model for contact analysis of the herringbone gears is established in ABAQUS. all the A tooth is divided into six parts to mesh with the grid, and then the grid of the tooth flanks is refined with a minimum size of 0.05 mm. To improve computational efficiency, the C3D8R element is defined to mesh the gears, as shown in Fig. 2. In addition, section property is assumed to be isotropic, and the mechanical parameters of the gear material are listed in Table 2 [8]. The input torque is set as 408 N\u22c5m, and the interaction is defined as surface-to-surface contact. The master surfaces are set as the pinion flanks and the slave surfaces are set as the gear flanks. The friction factor is set as 0.05, which is approximately the average value under oil lubrication [31]. And the friction formulation is set as penalty. To obtain continuous stress distribution results, a multi-load step method [32] is applied to the finite element analysis (FEA) model, as follows: (1) applying a minor torque in the z-axis direction to the driving wheel; (2) gradually increasing the driving gear torque to the rated torque and (3) applying rotation displacement to the driving gear, limiting the maximum increment size for each step to obtain a continuous meshing process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001752_j.mechmachtheory.2020.103992-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001752_j.mechmachtheory.2020.103992-Figure4-1.png",
        "caption": "Fig. 4. Oblique coordinate system S \u03b1 .",
        "texts": [
            " 3 , n can be implicitly expressed by cos (\u03c0/ 2 \u2212 \u03b1n ) = < n , r (1 rad) >, (15) where r (1 rad) [ 1 ] = ( sin (t) cos (t) 0 )T . (16) Further, n is orthogonal with the tangent direction of (1) , which means n r (1) \u2032 = 0 , (17) and n is a normalized vector; therefore, n n = 1 . (18) Eqs. (15) , (17) , and (18) are independent, and n is determined by the three independent equations because n is an R 3 vector. To obtain the expression of the normal vector from \u03b1n , an oblique coordinate system S \u03b1 is built at the point of (1) as shown in Fig. 4 . Let the covariant basis vectors of S \u03b1 be r (1) \u2032 , r (1 rad ) , and r (1) \u2032 \u00d7 r (1 rad ) . Subsequently, synthesizing Eqs. (2) and (16) in S 1 , the three covariant basis vectors of S \u03b1 can be represented as g ( S \u03b1 ) 1 [ 1 ] = d r (1) [ 1 ] (t) /dt = ( p sin (t ) f \u2032 (t ) + p f (t ) cos (t ) , p cos (t ) f \u2032 (t ) \u2212 p f (t ) sin (t ) , f \u2032 (t ) )T g ( S \u03b1 ) 2 [ 1 ] = r (1 rad) [ 1 ] = ( sin (t) , cos (t) , 0 ) T g ( S \u03b1 ) 3 [ 1 ] = g ( S \u03b1 ) 1 [ 1 ] \u00d7 g ( S \u03b1 ) 2 [ 1 ] = ( \u2212 cos (t) f \u2032 (t) , sin (t) f \u2032 (t) , p f (t) )T "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001632_j.ijrefrig.2020.03.029-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001632_j.ijrefrig.2020.03.029-Figure4-1.png",
        "caption": "Fig. 4. Structure comparison of conventional compressor and PSF centrifugal compressor.",
        "texts": [
            " The 60 \u00b0 circular refrigerant injection port not only enhances the flow niformity at the entrance of the second stage impeller, but also educes the influence of gas turbulence on impeller efficiency. .3. Two-stage impeller directly driven by PSF motor Mechanical efficiency is an important factor to evaluate the entrifugal compressor performance. The conventional compressor ot only adopts a pair of speed-increasing gears to meet the equirements of high pressure ratio, but also needs 4 supporting oints of radial bearings due to the motor speed limitation, as hown in Fig. 4 (a). The mechanical loss of speed-increasing gears nd 4 radial bearings accounts for about 4% of the power conumption under full-load condition. Under part-load condition, he mechanical loss can reach 8% of the unit power consumption. hat\u2019s more, the gear pair moves relatively in the high-speed perating process, which produces the transfer torque and a strong ibration. PSF centrifugal compressor improves the conventional transission mechanism based on high speed permanent-magnetic ynchronous motor. The speed-increasing gears for conventional ompressor are removed. The single axis and 2 radial bearings re adopted to achieve power transmission and structural stability, s shown in Fig. 4 (b). The improved structure has the following dvantages: ( 1 ) the motion parts of centrifugal compressor are educed and the structure is more simple and reliable; ( 2 ) the adial load of bearings and axle weight decrease greatly, and the tability and reliability of the compressor are enhanced; ( 3 ) mehanical loss just occurs in 2 radial bearings, and correspondingly he mechanical efficiency of compressor is greatly improved; ( 4 ) he speed-increasing gears are removed, so the size and weight f whole unit are greatly reduced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000166_j.robot.2019.103326-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000166_j.robot.2019.103326-Figure1-1.png",
        "caption": "Figure 1: Prototype of the Intelligent Cane (ver. 5). This cane robot has been developed for gait training to provide safer and more efficient gait training for elderly and physically challenged people.",
        "texts": [
            " The rest of this paper is organized as follows: Section 2 presents the mechanism and the control architecture of our220 cane robot. Section 3 introduces indices for physiological cost evaluation. Section 4 presents a pilot experiment to clarify the relationship between the exercise load with an assistive device and the physiological effects for the users. Section 5 presents a clinical experiment to evaluate the225 feasibility of the proposed assistive strategies with the cane robot. Finally, Section 6 concludes the paper and presents future work. 2.1. Design and specifications of the cane robot230 Figure 1 shows a prototype of the Intelligent Cane (ver. 5)1 developed in our laboratory [28]. This cane robot is designed to help the elderly and patients walk and provide them with effective gait training. This robot consists of an admittance controlled omni-directional mobile235 base, a touch monitor, a controller box with a miniPC, and a sensor system including a 6-axis force/torque sensor, a laser range finger (LRF), an RGB-D camera, and an ear-clip pulse oximeter. One of the notable features of this robot is that the movement of the mobile base can240 be determined according to the user\u2019s intention by the applied force to the handle gripper and the user\u2019s posture measured by the LRF",
            " prevent fall-over by increasing the mass and damping parameters of the admittance controller [29]. In the low-level controller, the velocity of the robot is controlled according to the desired behavior of the robot310 determined in the high-level supervisor. An appropriate walking load provided to the user is selected by adjusting the virtual mass and damping coefficients in the admittance control model. The base motion of the cane robot has three degrees315 of freedom consisting of two translational motions and one rotational motion. As shown in Fig. 1, the translational motions are defined by the motion along the horizontal axes (x-axis and y-axis), and the rotational motion is defined as the rotation around the vertical axis (z-axis).320 The velocities of the cane robot are given by the following admittance control model: Mvv\u0307 + Dvv = f , (1) where v = [vx vy \u03c9z]T is the vector composed of the horizontal velocities and the angular velocity of the cane robot, f = [fhx fhy \u03c4hz]T is the vector composed of the horizontal forces and the torque around the z-axis applied to the cane robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003679_s0167-8922(08)70490-5-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003679_s0167-8922(08)70490-5-Figure1-1.png",
        "caption": "Fig. 1 Finite Element model of a main be a r i n g",
        "texts": [
            " Journal temperature The journal temperature is assumed constant and has to be supplied to the calculation procedure. If an estimation is used for the calculation, it can be checked by evaluating the amount of dissipation energy leaving the bearing through the journal. Because of symmetry considerations only a small fraction of up to 10% of the total heat generated is expected to be conducted through the journal (and the whole crankshaft). 2.6. Elasticity equation The consideration of the deformation of the bearing shell due to the oil film pressure requires a Finite Element model of the bearing shell (Fig. 1). 654 T h s model is reduced to the radial degrees of freedom of the nodes lying a t the inner surface by static condensation. With the resulting mass, damping and stiffness matrices the elasticity equation for the bearing shell is M i i , + D u , + K u , = f . This equation is solved by means of an implicit time integration method. 2.7. Equation of motion of the journa l The journal is modeled as a single mass yielding as equation of motion bearing shell and the calculation of the elastic deformations. Typically, 120 to 180 nodes in circumferential direction and about 25 nodes in axial direction are employed for the discretization of the Reynolds equation and the energy equations. Radially 7 to 14 nodes are used for the energy equation of the oil film and 7 for the energy equation of the shell. In the finite element meshes of the shell surface there are a t least 40 nodes circumferentially and 3 nodes axially. The model depicted in Fig. 1 possesses 48 x 9 nodes. For details about the solution process refer to [6,8]. m S,, = f, + f, , the equilibrium of inertia, oil film and outer forces. The latter are an input to the calculation procedure. 2.8. Viscosity Oil viscosity is regarded as a function of both local oil pressure and local oil temperature following Rodermunds formula For the parameters A through F of an SAE low-40 oil see Table 1. 2.9. Density For the dependence of oil density on pressure and temperature the DowsonHigginson formula, f, 'P 1 + f , ' P P( P1 T ) = Po ",
            " The minimum oil temperature is 657 always the temperature of the oil in the oil groove (97-deg-C). This results in a maximum difference of oil temperatures in the bearing of 29.5 deg-C (Fig. 7, Fig. 8). The maximum temperature rise from the inlet temperature of 90 deg-C is 36.5 deg-C. On the shell surface there is a temperature variation of 28 deg-C, with the minimum again a t the oil groove, Fig. 9. 3.2. Main bearing of a gasoline engine The Finite Element structure of this bearing is depicted in Fig. 1. Peak oil film pressure. (Fig. 10) Comparing the peak oil film pressure for pressure and temperature dependent viscosity and density ( q = q(p,T)) with the pressure for constant viscosity and density ( q = ~ lp=lbar ,T=120deg-c ) one can see that the absolute maximum over the engine cycle hardly changes for the two variants. Nevertheless, significant differences in the peak oil film pressure a t 700-60, 165- 180, 310-385 and 500-550 deg-CA can be observed. The maximum relative difference is almost 100% a t 20 deg-CA"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000168_3343055.3359704-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000168_3343055.3359704-Figure2-1.png",
        "caption": "Figure 2. Geometrical Representation of Bessel Beam: left) An elongated and slim sound beam (i.e., Bessel beam) is created as a result of a conical and propagated wavefronts, and right) information needed to create the illustrated wavefront from a transducer in a Phased Array Transducer (PAT).",
        "texts": [
            " Basic Bessel beam creation In this section we explain our technique, which is compatible with both PATs and hybrid modulators, and uses Bessel beams to create a stream of air particles (i.e., move the aerosol), redirect it in real-time (i.e., allow mid-air control), and to retain the laminarity of the flow (i.e., for display quality purposes). We refer to a related work by Hasegawa et al. [7] to construct the Bessel beam. In general, one can produce this Bessel beam with a conical arrangement of sound sources, where the sound waves will converge and concentrate its energy (i.e., ultrasonic radiation force) as in Figure 2(left). To construct the beam, we use a modulator (PAT or MM), operating at an ultrasonic frequency ( f = 40kHz) in air (speed of sound c = 343ms\u22121). The algorithm includes two steps (see Figure 2(right)): Firstly, we compute the angle \u03b8z in Equation (1), given a constant aperture A (modulator diameter), and zm as the maximum height of the beam. \u03b8z = tan\u22121 ( A 2 \u00b7 zm ) (1) Secondly, we compute the phase for each element in our modulator as in Equation (2), where k = 2\u03c0/\u03bb is the ultrasound wavenumber, and \u03bb = c/ f is the wavelength, in our case 8.6mm, and d(Ti,0) is Euclidean distance function. The phase profile \u03c6i of the element can be electronically delivered to a PAT, or fabricated into a MM as in [18]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002542_icem49940.2020.9270965-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002542_icem49940.2020.9270965-Figure8-1.png",
        "caption": "Fig. 8. Schematic of axial and circumferential segmentation [23].",
        "texts": [
            " A modular stator PM synchronous machine has been proposed which applied for wheel-driving vehicle application as well [26], shown in Fig. 7. The 24 slots 14 poles combination, which exhibited the harmonics inhibition of magnetomotive force (MMF). The highlight of topology, concentrated winding configuration with unequal teeth widths, has completely isolation of thermal and magnetically. For specific aerospace actuator applications, two candidate segmentation strategies, axial segmentation and circumferential segmentation are adopted for fault tolerance and shown in Fig. 8. [23]. In the formal architecture, two separate stators with separate driven inverters are embedded into a common housing and rotor, adopting single layer FSCW configuration; the latter one, two set of three-phase in one stator with double-layer FSCW configuration driven by two independent inverters. The outputs of simulation and experimental tests showed the excellent fault tolerance. To ensure adequate levels of functional safety, a FT concept for the design of in-wheel motors has been demonstrated [25], in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002914_j.cirp.2021.04.063-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002914_j.cirp.2021.04.063-Figure1-1.png",
        "caption": "Fig. 1. Shape and scanning method of AMed metal workpiece.",
        "texts": [
            " However, aluminum is well known as an active metal, and the surface is easily oxidized. Thus, it is difficult to prevent the oxidation of aluminum alloy surface in PBF process in reduction pressure atmosphere. Then, the surface characteristics of AMed aluminum alloy may deteriorate. Therefore, in this study, surface smoothing and repairing of AMed maraging steel and aluminum alloy products by large-area EB irradiation were experimentally investigated. The change in surface elemental composition was also discussed. Fig. 1 shows the shape of AMed metal workpiece and scanning pattern in PBF process. PBF conditions are shown in Table 1. A commercial PBF machine (ProX DMP 200, 3D systems) was used to fabricate the workpiece. Laser scanning pattern was alternated with orthogonal directions (90\u00b0) from layer to layer, since the laser scanning direction is typically changed in each layer to obtain more uniform AMed products [10,11]. Then, the workpiece with 10\u00a3 10\u00a3 5mm was produced. Maraging steel (Fe-18wt%Ni10wt%Co-5wt%Mo) and Al-Si aluminum alloy (Al-12wt%Si) powders were used in PBF process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002218_j.mechmachtheory.2019.103771-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002218_j.mechmachtheory.2019.103771-Figure3-1.png",
        "caption": "Fig. 3. A novel parallel-link mechanism for forging manipulators.",
        "texts": [
            " Section 2 analyzes the principle of motion decoupling of a novel main-motion mechanism, and determines the dimension parameters to establish the virtual prototype. In Section 3 , the rigid-body dynamic and the elastodynamic models are constructed, respectively. Section 4 presents the solutions to the elastodynamic equations and the elastodynamic response curves of this novel forging manipulator. Conclusions are provided in Section 5 . In this section, a novel parallel-link mechanism for forging manipulators is proposed, and its main-motion mechanism is shown in Fig. 3 . The horizontal buffering mechanism is composed of an equivalent mechanism of the Hoeckens straight-line mechanism (see Fig. 4 ) when the buffering cylinder q 3 is locked, and the ratio of the length of each linkage satisfies l PN : l PH : l N G : l MG : l GH = 1:2:2.5:2.5:2.5 [24 , 25] . The output point ( M ) of the Hoeckens straight-line mechanism is directly connected to the middle part of the gripper carrier ( EF ), and the buffering cylinder q 3 is installed into the linkage HG of the straight-line mechanism to form a horizontal buffering mechanism"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure22.4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure22.4-1.png",
        "caption": "Fig. 22.4 Exoskeleton analytical model: a at the beginning of the load lifting process; b at the end of the load lifting process",
        "texts": [
            " The position of the mechanism links is determined by absolute angles \u03c6i . The most important issue during exoskeleton modeling is the kinematical and force parameters of the device movement determination that allows to achieve the required load movement. Thus, in the framework of this study, the modeling task consists in determining the time dependences of the exoskeleton links rotation angles during the load movement along the given trajectory, as well as determining of the joints torques. Let us consider analytical models shown in Fig. 22.4. A number of assumptions were accepted in the study of load lifting process: \u2022 The load lifting process is carried out along a straight vertical line that means that during the lifting process, the load securing pointO4 changes only one coordinate y04. In this case, the law of the O4 point coordinates change i, represented by the radius vector r04 in time, can be represented in a polynomial form: r04 = ( 0, 5\u2211 i=5 ki t i )T . (22.1) \u2022 The cargo lifting process is realized with a fixed angle of the exoskeleton\u2019s back inclination \u03c63 = const = 80\u25e6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001438_lra.2020.2969161-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001438_lra.2020.2969161-Figure10-1.png",
        "caption": "Fig. 10. User test. (a) Setup consisting of the forceps-driver equipped on the tele-operated microsurgical robotic system and the master device installed on a 6-DOF haptic device (Phantom Omni, Sensable Technologies). (b) The stiffness variation of master device varied according to the input displacement.",
        "texts": [
            "28 ms, and the latency from the gripping-force detection to stiffness variation is measured as about 0.1 ms. The experimental results for the repetitive operation are indicated in Fig. 9(b-1) and (b-2). The experimental result shows that the stiffness variation pattern is not in directly proportional to the gripping-force variation pattern. As a kind of control issue, this could be improved by applying the high-precision control scheme for the SMA [16]. To conduct the user test, the prototyped devices are installed on the previously implemented microsurgical robotic system. As shown in Fig. 10(a), the forceps-driver is mounted on the end of the robotic manipulator as an end-effector, and the master device is assembled in a 6-DOF haptic device (Phantom Omni, Sensable Technologies). By utilizing this setup, the user performs the task to tele-operate the forceps-driver and grip the object. In the experiment, the input-output relationship is evaluated quantitatively as the results plotted in Fig. 10(b) first, and then the user test is performed to survey the user satisfaction. The users who attend this test report that they can feel the different haptic feedback according to the material of the object sample. However, they also point out some critical drawbacks of the proposed system. The first is that the size of the prototypes master device is quite large to hold and operate the device. Because the size of the device can be easily reduced by designing the small-size diaphragm flexure, the master device will be implemented as a compact and lightweight device through the modification of the form factor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002078_fie44824.2020.9274250-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002078_fie44824.2020.9274250-Figure1-1.png",
        "caption": "Fig. 1. An example of decomposition item question in the ECTD",
        "texts": [],
        "surrounding_texts": [
            "2. How do the learning gains of students differ by gender and race? III. METHOD Participants. The population of the study was the entire 2019 first-year engineering cohort at a large Southwestern university in the United States. Institutional Review Board (IRB) approval was obtained and participants were recruited via IRB-approved emails. To avoid participation bias, recruitment emails were distributed to all first-year engineering students by faculty members not affiliated with the research team. Pre-test recruitment emails were sent twice in the second week of the Fall semester. About 800 students responded to the pre-test version of the ECTD. After factor analysis to gauge the effectiveness of the ECTD questions, changes to questions were made before post-test use of the ECTD. At the start of the Spring semester, two recruitment emails were sent to recruit participants from the same population for post-test measurement. While around 100 students responded to the post-test version of the ECTD, 62 students were identified for completion of both the pre-post ECTD. Table II shows demographic information of these 62 participants. TABLE II. PARTICIPANTS\u2019 DEMOGRAPHICS Gender Race and Ethnicity Sub-total Hispanic AIAN Asian Black Multiracial White Female 1 1 3 0 3 10 18 Male 7 0 10 1 0 26 44 Total 8 1 13 1 3 26 62 Note. AIAN = American Indian or Alaska Native Authorized licensed use limited to: Konya Teknik Universitesi. Downloaded on May 20,2021 at 08:16:06 UTC from IEEE Xplore. Restrictions apply. Data Analysis. Based on the sample size, the research team conducted a paired sample t-test or Wilcoxon signed ranks test, which is a counterpart of the paired sample t-test in nonparametric tests, to determine if there is a significant difference between the pre- and post-ECTD results. Due to the small sample size by subgroups of gender and minority status, we conducted a Mann-Whitney U test, which is a counterpart of independent samples t-test in nonparametric tests, to explore subgroup differences in the pre-post changes on the ECTD scores. Here, minority status was grouped as White versus nonWhite students due to the small sample size. IV. RESULTS Table III shows the descriptive statistics on the pre-post ECTD scores by gender and minority status. The differences between the pre-post mean scores were all positive, indicating the gain of student learning on computational thinking. Table IV shows the statistical testing results on the prepost differences by gender and minority status. The paired sample t-test indicates a significant mean difference of 2.26 between pre and post-ECTD scores with t(61) \u2013 7.4, p < 0.001 with a significant correlation of r = 0.480 between pre-post scores. Similarly, paired sample t-test and Wilcoxon signed ranks tests shows significant increases on the post-scores for each subgroup by gender and minority status. The Cohen\u2019s d effect sizes on the pre-post changes were all positive and medium to large [38]. Perceptually, the magnitudes of the change were larger for female and White students. male students increased 2.21 points on the post-ECTD. However, the Mann-Whitney U test shows nonsignificant difference by gender on the changes of the pre-post scores, with U = 301.0, Z = -1.489, p = 0.137. Similarly, non-White students increased 2.23 and White students increased 2.97 points on the post-ECTD. However, the Mann-Whitney U test shows nonsignificant difference by gender on the changes of the pre-post scores, with U = 375.5, Z = -1.332, p = 0.183."
        ]
    },
    {
        "image_filename": "designv11_14_0000876_0954407018824943-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000876_0954407018824943-Figure13-1.png",
        "caption": "Figure 13. Neighbouring and wrapped belt-spans of a generic driven pulley i.",
        "texts": [
            "16 The non-linearity due to potential periodical stick slip motion of the tensioner22 is not considered in the present analysis. The objective is to determine the periodic rotation angle of the tensionerarm and the periodic fluctuations of the belt-span tensions. These last quantities are the inputs for the model that estimates the hysteresis losses in the tensioner and in the belt-spans. The tension fluctuations DT (e.g. Figure 12) are the cause of additional belt-stretching hysteresis power losses in the free \u2018j\u2019 and wrapped \u2018i\u2019 belt-spans (Figure 13). The belt-stretching hysteresis power losses due to the difference between the equilibrium belt-span tensions17 can be calculated through equation (5) as performed by Silva et al.11 Similarly, the additional energy Wspanj and power PLspanj losses by belt-hysteresis due to the variation of belt free span tensions, for example, DTj with j=1, 2, . . . , 6 for the FEAD in Figure 12 can be calculated by equations (36) and (37), respectively Wspanj =pE00Dej2 \u00f036\u00de PLspanj = lWspanjV \u00f037\u00de where E00 is the loss modulus of the belt-elastomer obtained via DMA; V, the belt linear velocity; Dej, the variation of strain to which the belt is subject during the belt travel along the free span \u2018j\u2019 (Figure 13); and l, a constant dependent on the belt-span length, the engine speed and the frequency of the tension fluctuations. Indeed, the quantity WspanjV in equation (37) represents a complete loop of hysteresis (Figure 2); however, if along a generic belt-span, the corresponding harmonic tension (Figure 12) has time to complete two loading cycles, then l=2 in equation (37). At the opposite, if the period of tension fluctuation is less than the time required to travel the belt span length, then 04 l4 1. Moreover, according to Hooke\u2019s law, that is, equation (23), the variations in belt-span tension, generated from the engine acyclism, leads to variation in strain Dej and vice versa, equation (38) Dej = DTj EA \u00f038\u00de where DTj is the tension variation in a free belt-span j (or j \u2013 1 leading to DTj 1) (Figure 13) and EA the belt tensile modulus or strain stiffness along the belt. For the wrapped belt-span \u2018i\u2019 in Figure 13, the energy Wspani and the power PLspani losses as well as the variation in strain Dei are calculated exactly as equations (36)\u2013(38). However, here as in Barker et al.,20 the variation in tension DTi leading to Dei is taken to be an average, so that the DTi acting on the wrapped belt span \u2018i\u2019 in Figure 13 is the sum of the neighbouring belt-span tension variations DTj and DTj 1 divided by 2, equation (39) DTi = DTj +DTj 1 2 \u00f039\u00de The total power loss by belt-vibration PLvib can be calculated by equation (40) PLvib = Xnp i= j=1 PLspani +PLspanj \u00f040\u00de Alternatively, PLvib can also be written as a function of tension fluctuations DTj, equation (41), by introducing equations (36)\u2013(39) into equation (40). DTj are obtained via a dynamic model16 PLvib= Xnp j=1 pE00 EA\u00f0 \u00de2 5 4 DTj 2 1 2 DTjDTj 1+ 1 4 DTj 1 2 V \u00f041\u00de Note that both ej and ei are cyclic since they depend on the cyclic belt-span tensions Tj and Ti leading to small belt-stretching hysteresis loops (Figure 2). Moreover, these losses can be directly associated with the damping (dissipation) inside the belt-spans in Figure 13. Also, there are as many free belt-spans as there are wrapped belt-spans in a FEAD. For example, applying equations (40) and (41) to the FEAD in Figure 11, \u2018j\u2019 and naturally \u2018i\u2019 would range from 1 to 6. In addition, attention must also be paid to the first pulley during the analysis. In this case, the j 1 in equations (40) and (41) shall be replaced by np, where np is the number of pulleys composing the FEAD. Simulation and experimental results are presented for the serpentine belt drive system defined in Table 1 and presented in Figure 14(a), where the CS pulley is assimilated to the driving pulley, the two others AD1 and AD2 are accessory driven pulleys"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002417_j.isatra.2020.07.042-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002417_j.isatra.2020.07.042-Figure1-1.png",
        "caption": "Fig. 1. Interception guidance sketch.",
        "texts": [
            " Usually, we find that 1 should be chosen as a vector that has a large norm and F2 is sually selected as a matrix that has a small eigenvalue. The NDO ain p will trade off the convergence rate and observation error f the proposed NDO. . Simulation .1. Missile\u2013target guidance problem To apply the presented adaptive optimal algorithm to the uidance problem, the missile\u2013target dynamics is built firstly. Consider a planar engagement scenario. \u03b1 is flight path angles FPA) of interception missile M , \u03b2 denotes the FPA of target T , hown as in Fig. 1. \u03b8 is defined as line of sight (LOS) angle. ts angular rate along the LOS \u03b8\u0307 can be designated as \u03c3 . Their relative distance is r . Vr denotes the range rate along the LOS. Besides, VM , VT and uM , vT are their velocities and control inputs, respectively. The relative dynamics is described as follows. Vr = VT cos(\u03b2 \u2212 \u03b8 ) \u2212 VM cos(\u03b1 \u2212 \u03b8 ) r\u03c3 = VT sin(\u03b2 \u2212 \u03b8 ) \u2212 VM sin(\u03b1 \u2212 \u03b8 ) \u03b1\u0307 = uM/VM \u03b2\u0307 = vT/VT (72) Inspired by the work of [41], if we adjust the variable \u03c3 to zero while keep the variable Vr to be negative, then r \u2192 0 eventually, which indicates this interception task is finished"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001459_s42835-020-00350-8-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001459_s42835-020-00350-8-Figure1-1.png",
        "caption": "Fig. 1 Motions of quadrotor UAV",
        "texts": [
            " The simulation results demonstrate that the proposed control system when the adaptive fuzzy controllers exist in the first layer has superiority over the control system when all the three layers have TPIDCs only. In addition, the obtained results from GS illustrate that the proposed control system based on the PSO algorithm managed to drive the QUAV on the desired path feasibly and smoothly. Moreover, the VFHA enhances the ability of the QUAV to avoid hitting the barriers blocking the desired path. 1 3 The quadrotor UAV has four propellers to generate suitable forces ( F1,F2,F3 andF4 ) for taking-off, following a trajectory and landing purposes as shown in Fig.\u00a01. The four propellers can be considered as two pairs: the propeller1 and propeller3 are the first pair while propeller2 and propeller4 are the second pair. In order to balance the torques and make the QUAV stable during the maneuvering, the first pair should rotate clockwise while the second pair should rotate counterclockwise or vice versa. By changing the related torque between each pair, different motion types can be generated such as yaw motion, pitch motion, roll motion, and altitude movement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000748_978-981-10-2875-5_114-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000748_978-981-10-2875-5_114-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of the pneumatic muscle actuator",
        "texts": [
            " In consideration of this issue, this paper deals with the motion control of PMA using fast switching valve. The rest of this paper is organized as follows. In Sect. 2 the experimental system is briefly introduced. In Sect. 3, the static model, dynamic model, and hysteresis model of the PMA are systematically derived. Then, a close-loop control scheme is proposed to achieve high accuracy trajectory tracking control of the PMA, and the simulation is carried out in the environment of MATLAB/Simulink in Sect. 4 before conclusions are drawn in Sect. 5. Figure 1 shows the schematic diagram of the pneumatic muscle actuator. The components used here are given in Table 1. The air compressor connects the fast switching valve with throttle valve and reservoir. The fast switching valve 1 is inlet valve, called inlet valve for short; the fast switching valve 2 is exhaust valve, called exhaust valve for short. Without the excitation of the solenoid, the valve will keep closed due to return spring. The working process is as follows. Initially, the PMA connects with the external environment, and the internal pressure of PMA is equal to atmospheric pressure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002355_tec.2020.3001914-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002355_tec.2020.3001914-Figure2-1.png",
        "caption": "Fig. 2. Rotor assembly of the PMSM under study.",
        "texts": [
            " Specifically, we consider the effects of uncertainty on average torque, sixth torque harmonic magnitude, peak line-to-line voltage, and total core loss. The device under study is an eight-pole PMSM, rated for 235 Nm and 96 kW at 3900 rpm. Its cross-section is depicted in Fig. 1, where \u03b8rm is the mechanical rotor angle. The rotor lamination stack is divided into four identical sub-stacks, such that the rotor stack length equals the stator stack length, `m = 90 mm. During assembly, two of the four sub-stacks are rotated clockwise, whereas the other two are rotated counterclockwise, as shown in Fig. 2. A skew angle, \u03b8sk, is defined to parameterize this rotation. Note that Fig. 1 shows the cross-section of a sub-stack with a positive skew angle. It should be noted that the qd-axes are representative of an average magnetization direction from the combined effect of all four sub-stacks. To analyze sub-stacks, first we develop a parametric two-dimensional FEM. This high-fidelity FEM provides input-output observations that are used to construct a computationally inexpensive surrogate model that eventually replaces the FEM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001752_j.mechmachtheory.2020.103992-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001752_j.mechmachtheory.2020.103992-Figure11-1.png",
        "caption": "Fig. 11. Solid model of the pinion and gear.",
        "texts": [
            " (80) To avoid a case where the tooth root thickness of the pinion or the gear is too thin, the arc radius of the cross-section curves of the pinion and the gear are set to be the same magnitude, which means R (1) = R (2) . From (80) , the arc radius can be can calculated and rounded off as R (1) = R (2) = 25.8 mm. Similarly, the cross-section curves of the convex side can be set to be the same as those of the concave side. In the convex side, K (12 u ) is set as 0.236 mm \u22121 , and R (1) = R (2) = 8.4 mm. Finally, from Eqs. (8) \u2013(12) , the tooth surface of the pinion and the gear can be exported, and the solid models of the pinion and the gear can be built as shown in Fig. 11 . 5.5. Stress analysis Based on finite element analysis, a loaded contact analysis can be completed to evaluate the stress distributions of the pinion and gear in Fig. 11 . The finite element model is as shown in Fig. 12 , which is meshed by the first-order hexahedral element. For the elements on the tooth surface, the max size of element length is set as 0.2 mm, and for other elements, the max size is set as 2 mm. The Young\u2019s modulus is 2.1 \u00d7 10 11 Pa and Poisson\u2019s ratio is 0.267. The torque of 540 Nm applied on the gear and the pinion is fixed as shown in Fig. 13 . The results of finite element analysis are shown in Fig. 14 . It can be seen that the shape and orientation of the contact ellipse of the finite element analysis are similar to those of the analytical analysis in Section 5"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003800_0043-1648(95)06911-9-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003800_0043-1648(95)06911-9-Figure2-1.png",
        "caption": "Fig, 2, Rotating lldn disc geometry and boundary conditions,",
        "texts": [
            " For the exact definiqon of h, knowledge about the surface and time factors, the geometry, streamlined conditions and the physical properties is needed. The most important quantities for physical applications are the maximum and the average temperatures over the area of nominal contact. As in the investigations [8-10], we used the average steady temperature as the basis for the calculation of the friction heat distribution coefficient. 3. Heat transfer model for the rotating disc The geometry and boundary conditions are shown in Fig. 2. It is assumed that the coordinate system (r, O) is fixed to the heat source and the disc rotates with respect to this coordinate system with the speed ca. Taking the assumptions ( 1 )-(6) into account, the governing differential equation in dimensionless form is [ 11 ] 02T IOT h ' _ ~ST ~ + ; ~ - ~ r = ~ (I) subjected to the boundary condition Ka.~=fq lOl<Oor=R r L-hT lO]>Oo=R (2) Using the dimensionless variables we rewrite the Eqs. (1) and (2) as follows: 82T * I OT* . OT* OT* ~1 I OI < oo p= 1 \"~'p=L-BiT* 10i>0o p=l (4) Using the finite Fourier transform of the form [ 12] 11' E 7\"*(p,n) =~-~ T*(O, 0)e-i\"\u00b0d0 - 'n - where 1 n--O e= 2 n~O A, Yevtushenko el aL / Wear 196 f1996 ) 219--225 221 with the regularity condition T*(p, -'rr)-T*(p, ~) the Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002020_s12206-020-1030-6-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002020_s12206-020-1030-6-Figure4-1.png",
        "caption": "Fig. 4. Extreme posture of finger mechanism.",
        "texts": [
            " l l l l l l \u03b8 \u03b8 \u03b8 \u03b8 + + + + = (11) For given 1l , 2l , 3l , 4l and 1\u03b8 , 3\u03b8 could be obtained by solving the above nonlinear trigonometric equation. The displacement of point F could be obtained as F B BF= +R R R (12) where, 3 3 5 3 3 3 5 3 cos sin 2 sin cos . 2 BF l l l l \u03b8 \u03b8 \u03b8 \u03b8 \u239b \u239e= \u2212 +\u239c \u239f \u239d \u23a0 \u239b \u239e+\u239c \u239f \u239d \u23a0 R i j (13) By substituting Eqs. (1) and (13) into Eq. (12), FR can be obtained as 4 3 1 1 3 5 3 3 1 1 3 5 3 cos cos sin 2 2 sin sin cos . 2 F l ll l ll l \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u239b \u239e= + + \u2212 +\u239c \u239f \u239d \u23a0 \u239b \u239e+ +\u239c \u239f \u239d \u23a0 R i j (14) As depicted in Fig. 4, the initial value of 1\u03b8 can be obtained by assume 3\u03b8 = 0 as ( )1, 4arccos 2s AKl l\u03b8 = \u2212 (15) where, AKl denotes the distance between points A and K in the initial state. Point K denote the intersection point of link AB and CD. At the initial state, AKl could be obtained directly according to the geometric analysis as ( )1 4 3 4 .AKl l l l l= + (16) Combine Eqs. (15) and (16), 1,s\u03b8 could be rewritten as ( )( )1, 3 4 1arccos 2 .s l l l\u03b8 = + \u2212 (17) As finger moves to the extreme state (as depicted in Fig. 4), the extreme value of 1\u03b8 could be obtained by cosine formula as 2 2 2 2 2 3 2 3 1 4 1, 1 4 2arccos . 2e l l l l l l l l \u03b8 \u239b \u239e+ \u2212 \u2212 \u2212= \u239c \u239f \u239d \u23a0 (18) Here, we define an average transmission ratio rT and operation range parameter r\u03b8 , to evaluate the kinematic behavior of the finger mechanism 3, 1,=r r rT \u03b8 \u03b8 (19) 3, 1,=r r r\u03b8 \u03b8 \u03b8+ (20) where, 3, 3, 3,=r e s\u03b8 \u03b8 \u03b8\u2212 and 1, 1, 1,=r e s\u03b8 \u03b8 \u03b8\u2212 represent the rotation range of link AB and BC. 3,s\u03b8 and 3,e\u03b8 denote the initial and extreme angle of 3\u03b8 , respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002362_tie.2020.3000088-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002362_tie.2020.3000088-Figure10-1.png",
        "caption": "Fig. 10. PWM voltage injection using the proposed PWM when the position is at (a) \u03b8e=0-deg, (b) 60-deg, (c) 90-deg and (d) 210-deg",
        "texts": [
            " 9(b) using only these three active vectors and the zero vector (0,0,0). It is noteworthy that the DC bus usage is decreased without other three active voltage vectors. In addition, the proposed PWM might result in a little higher EMI noises because the zero vector (1,1,1) cannot be applied. Fortunately, the saliency-based drive is only applied for low speed operation where the EMF voltage is insufficient to estimate. Besides, EMI noises can also be reduced by cable grounding, wiring and shielding. Fig. 10 illustrates the switching frequency d-axis squarewave voltage injection at four different rotor positions using the proposed PWM in Fig. 9(b). At Fig. 10(a) \u03b8e=0deg, the negative square-wave voltage is generated through the combination of (0,1,0) and (0,0,1). Similar combination can be used for \u03b8e= 60deg in Fig. 10(b). In addition, for \u03b8e= 90deg in Fig. 10(c) and \u03b8e= 210deg in (d), both the positive and negative square-wave voltage are resultant using two of these three active vectors. Authorized licensed use limited to: Murdoch University. Downloaded on June 16,2020 at 08:48:24 UTC from IEEE Xplore. Restrictions apply. 0278-0046 (c) 2020 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. Fig. 11 illustrates A-phase current and three-phase PWM voltage waveforms using the proposed PWM pattern for the switching frequency voltage injection",
            " However, the asymmetric current waveform is observed because only three active vectors are used. It is noted that the zero vector (0,0,0) is used for the zero-voltage control at the bottom of PWM period. Thus, three-phase currents can be obtained based on the shunt current sensing. In addition, either one of three active voltages is applied at the peak of PWM period. At this time, two of three phase currents can be measured for the third phase current reconstruction by using the proposed PWM. On the other hand, for \u03b8e =90-deg in Fig. 10(c), both positive and negative square-wave voltages are resultant by selecting two of three active vectors (1,0,0), (0,1,0) and (0,0,1). Fig. 11(b) shows corresponding A-phase current and three-phase PWM voltages. Because the active voltage (1,0,0) is used for both positive and negative injection voltage, an active vector (1,0,0) appears in the peak PWM period. Under this effect, ib and ic are measured at this sample instant. The current difference can also be calculated for the position estimation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001442_ilt-10-2019-0435-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001442_ilt-10-2019-0435-Figure1-1.png",
        "caption": "Figure 1 Thrust bearing model configuration with spirally distributed textures",
        "texts": [
            " Then, the key geometry parameters for textures are optimized, with the maximum loading capacity and minimum friction force as the objectives. The basic approach is combining the multi-objective optimization method based on the response surface methodology with the computational fluid dynamics (CFD) technology, which can provide a fundamental design guide for textured thrust bearings. A fixed geometry eight-pad thrust bearing with spirally distributed rectangular textures is used as baselinemodel in this study. A total of 16 textures are uniformly distributed along the spiral line in the radial direction, as shown in Figure 1. Taking sinusoidal distributions as examples of spiral distributions in this study, the textured spiral distribution near the entrance edge is expressed as: d \u00bc Sa sin 2p=T R Ri Fr\u00f0 \u00de Fc (1) Then, the relationship between the spiral lines in the circumferential direction is controlled by Dc. In present study, the cycle number Nt represents the number of cycles between the length ofWt, which is expressed asWt/T. The smooth wall rotates at a constant speed, while the partially textured pad is assumed to be \u201cstationary.\u201d In the bearing, the lubricant flows in from the inner bearing surface and flows out from the outer bearing surface. Rotational periodicity is implemented, as shown in Figure 1(d). The parameters of bearing geometry and boundary conditions used in present study are listed in Table I. Through the mesh independent study, the divisions used across the film thickness and textured depth are all 6, and a 0.1 interval size is used in the other directions. The baseline model has approximately 1,180,000 grid elements. Multi-objective optimization method Qiang Li et al. Industrial Lubrication and Tribology The laminar flow model is used and the lubricant is treated as an incompressible Newtonian fluid with a constant viscosity and density"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001350_0954407019890481-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001350_0954407019890481-Figure13-1.png",
        "caption": "Figure 13 TMF damage of cylinder head: (a) temperature results at rated speed, (b) mechanical fatigue damage, (c) oxidation fatigue damage, and (d) creep fatigue damage. TMF: thermo-mechanical fatigue.",
        "texts": [
            " The submodel temperature and displacement at P0, P5, and P9 load conditions were measured by the static stress\u2013strain calculation of the global model, and the other thermal load and boundary conditions were calculated by the following equations T i\u00f0 \u00de=T5 T5 T9\u00f0 \u00de3K i\u00f0 \u00de \u00f03\u00de U i\u00f0 \u00de=U5 U5 U9\u00f0 \u00de3K i\u00f0 \u00de \u00f04\u00de where T(i) is the temperature of the submodel, U(i) is the displacement of the outer surface of the submodel, and K(i) is the temperature and displacement coefficient from test results. The temperature results at rated speed are shown in Figure 13. The TMF analysis was performed using the HEAT module in FEMFAT software. The sub-FE model contains all four cylinders of the four-cylinder engine structure. The fire deck of the four cylinders has been evaluated. By simulating the heating-up and coolingdown of the engine, the Jiangling Motor Company (JMC) low cycle fatigue (LCF) durability test specification is simulated. A transient nonlinear stress\u2013strain analysis was performed on the submodel. Four cycles were calculated in order to get a stabilized TMF cycle. As shown in Figure 13, oxidation fatigue damage plays the leading role in the total TMF damage of cylinder head. Figure 14 shows the calculated TMF lifetime for each subcylinder. The lowest TMF lifetime has been found in the intake\u2013exhaust valve bridge close to the exhaust valves of the second cylinder and amounts to 7096 cycles. No critical region in view of TMF has been identified on cylinder head fire deck. The lowest calculated values are above the required limit value of the JMC LCF durability test specification"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002042_j.addma.2020.101730-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002042_j.addma.2020.101730-Figure13-1.png",
        "caption": "Fig. 13. Process-induced meso-structure of the printed component with a sketched approx. crack growth direction (a) and the orientation of a propagated crack during testing (b).",
        "texts": [
            " Additive Manufacturing 36 (2020) 101730 plane and strand orientation are parallel) and 0\u25e6/90\u25e6 are quite similar, the orientation 0\u25e6 (crack plane and strand orientation are perpendicular to each other) shows a slightly different slope (m) and ordinate value (A). Since residual fatigue lifetime estimations are known to be quite sensitive to the exact value of A and especially m [40], these differences could significantly alter the final values of Nprop. To choose the right values of A and m, it is necessary to examine the local printing orientation at the position of the crack. Therefore, the Gcode of the component and the orientation of a propagated crack after testing were analysed (Fig. 13). When comparing these two, it is obvious that the crack initially has to grow perpendicular to the two perimeters (ergo 0\u25e6 orientation). Thus, the parameters A and m of the 0\u25e6 direction should be used. After the crack has propagated through the perimeter, the crack has to grow through an approximately \u00b1 30\u25e6/60\u25e6 strand orientation (as shown in Fig. 13a). However, this angle between the deposited strands and crack growth direction is primarily determined by the boundary conditions of the setup. In this case, mainly the geometry of the component and the friction between the PLA wrench and steel bolt. Due to the propagating crack, boundary conditions change continuously and the applied parameters A and m would have to change nonstop, for every increment of crack growth. Since it is not feasible to measure every possible combination of A and m, and the fracture surface (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001149_ab3e7a-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001149_ab3e7a-Figure2-1.png",
        "caption": "Figure 2. (a) 3D schematic of the laterally-moved tunable deflector. Two metasurfaces are cascaded with a gap d. The deflection angle \u03b8 is changed continuously by shifting the bottom layer metasurface. (b) Phase distributions of j1, j2 and their superposition jtot, after the motion of \u0394y=5176 nm, with a0=0.5\u00d710\u22124 nm\u22121.",
        "texts": [
            " We then calculated jtot of the cascaded unit cells in figure 1(b) by simulations corresponding to different nano-rods, and compared it with the sumption of the simulated j1 and j2 of single layer unit cells, which are shown in figures 1(c) and (d). It is found that they are very close. Figure 1(e) shows their difference jdif=|jtot\u2013(j1+j2)|. The maximum value is only \u03c0/7, which means jtot=j1+j2 can be applied for the selected gap distance. We then give the design of the deflector tuned by laterally displacing of the cascaded bilayer metasurfaces, which is shown in figure 2(a). The size of the first metasurface is \u22128.2 \u03bcm<x<8.2 \u03bcm and \u22124.2 \u03bcm<y<4.2 \u03bcm. The width along the x-axis of metasurface 2 is the same as metasurface 1 while that along the y-axis is 4 times larger than the former to avoid decreased effective incident aperture during the motion. It should be noted that the sizes of two layers can both be enlarged in practical fabrication [20]. Assuming the output ray is deflected to the x-axis, then the functional phase of a deflector is expressed as j q=x y k x, sin 20( ) ( ) ( ) in which \u03b8 is the deflection angle",
            " The phase distributions of the two metasurfaces are calculated by j =x y a k xy, 31 0 0( ) ( ) j = -x y a k xy, 42 0 0( ) ( ) where a0 is a constant determining the rate of phase variation. If metasurface 2 moves along the y-axis with \u0394y, the total phase retardation of the cascaded metasurfaces is j j j= + - D = Dx y x y x y y k xa y, , , . 5tot 1 2 0 0( ) ( ) ( ) ( ) Comparing equations (2) and (5), it is found that the relation between the deflection angle and the shifting distance is q = Da yarcsin . 60( ) ( ) Therefore, the deflection can be tuned by \u0394y if eligible a0 is used. To give an example, figure 2(b) shows the phase distributions of metasurfaces 1 and 2 on the left, corresponding to \u0394y=5176 nm, and a0=0.5\u00d710\u22124 nm\u22121. Their superposition is depicted on the right. The linear phase gradient along the x-axis indicates a tilted output beam according to the generalized Snell\u2019s law [27]. We then carry out the simulations. The simulation region is set as \u22128.2 \u03bcm<x<8.2 \u03bcm and \u22124.2 \u03bcm<y< 4.2 \u03bcm, which is the aperture of the device. The selection of the Si nano-rods is based on equations (3) and (4), as well as the results of figure 1(b)",
            " The decrease of the efficiency is attributed to two reasons. First, the working principle of a metasurface deflector is similar with a one-dimensional grating whose diffraction efficiency is decreased with the deflection angle [28]. Besides, with the increase of the deflection angle, less energy transmits the monitor in the simulation, which is located 1.5 \u03bcm away from the exit of the whole structure. Although it is possible to tune the beam deflection by drawing metasurface 2 along the y-axis in figure 2(a), the total size of the device cannot be miniaturized with the increased shift for larger deflection angles. To overcome this disadvantage, we designed another tunable deflector by twisting the metasurfaces like a Moir\u00e9 lens, which is shown in figure 4(a). The Si nano-rods are arranged in the polar coordinates. The total radius of the devices is 8.2 \u03bcm and d is 500 nm. The phase distributions of the two metasurfaces in the polar coordinates are constructed as: j a a=r k r, 1 2 sin 71 0( ) ( ) ( ) j a a= -r k r, 1 2 sin "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000234_ecce.2019.8913191-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000234_ecce.2019.8913191-Figure2-1.png",
        "caption": "Fig. 2. Rotor structures of the proposed VFMM and the target IPMSM. (b) The target motor",
        "texts": [
            " In this paper, the target maximum torque, physical parameters, and operating range were set as those of the target IPMSM used in the TOYOTA PRIUS fourth-generation. Table 1 lists target specifications and physical parameters of the proposed VFMM and the target IPMSM. The maximum current of the proposed VFMM is larger of 20 Arms than that of the target IPMSM. Instead of that, the dc-side voltage of the drive inverter is smaller of 60 V than the target IPMSM, and eventually, the inverter rated output capacity is same value of 132 kVA. Fig. 2 shows cross-sectional views of rotors employed by the proposed VFMM and the target IPMSM. In the proposed VFMM, VPMs are also installed in the rotor unlike the target IPMSM. In addition, the proposed VFMM has large flux barriers to enhance magnetization characteristic. The large flux barriers are called \u201cthe extended flux barriers\u201d in this paper. The proposed VFMM employs AC-magnetized structure because of advantage in terms of cost and same configuration as the target IPMSM. The both motors have same stator employing the segmented winding in order to increase the coil space factor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002830_s00170-021-06874-1-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002830_s00170-021-06874-1-Figure9-1.png",
        "caption": "Fig. 9 Five degree-of-freedom dynamic model of table system",
        "texts": [
            "99 \u03bcm/m respectively, and the roll error increases linearly under the action of vertical deviation. This shows that the table position only affects the roll error variation under the both horizontal and vertical deviations of one guide rail. The dynamic characteristics of table are affected by two harmonic forces (Fx sin \u03c9t and Fy sin \u03c9t) and three harmonic moments (Mx sin \u03c9t,My sin \u03c9t andMz sin \u03c9t). In order to investigate conveniently the vibration behaviors of table system, the table system is simplified to the mass-spring-damping model with five degrees-of-freedom as shown in Fig. 9, where \u03b4x and \u03b4y are the displacements of the geometric center of table relative to its initial position under the harmonic forces. The table deflects around the X-, Y- and Z-axes under the harmonic moments, which the deflection angles are \u03b5x, \u03b5y and \u03b5z, respectively. According to the simplified table system model, the dynamic equations can be expressed as follows M :: X\u00fe CX\u0307\u00fe Fk \u00bc F \u00f027\u00de where :: X andX\u0307 are the acceleration and velocity vectors, respectively;M and C are the mass matrix and damping matrix, respectively; Fk is the interaction force vector; F is the excitation force vector"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001732_15397734.2020.1776133-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001732_15397734.2020.1776133-Figure10-1.png",
        "caption": "Figure 10. Crushing mode of a 4-web corner element with angle 60 degrees between the webs with plastic deformation zones showing crushed state up to angle a:",
        "texts": [
            " By observing the top view, it is revealed that both half-waves of this 4-web configuration have the same crushing mode. This mode is shown in Figure 9d and according to plastic zones in the figure, the energy associated with each crushing wave of a regular 4-web corner element is stated as follows: X2 k\u00bc1 Ekn4:90\u00f0v, n\u00de \u00bc X2 k\u00bc1 2E1 p 4 \u00fe E2q p 4 \u00fe E3q p 4 \u00fe E2e p 4 \u00fe E3e p 4 \u00fe 2E4q p 4 k (8) There is another 4-web configuration found in the MT structures covered in this study which is shown in Figure 10. This one has two different crushing modes related to each half-wave as shown in Figure 10a,b and according to plastic zones shown in the figure, the energy associated with each crushing wave of this configuration is derived as follows: X2 k\u00bc1 Ekn4:60\u00f0v, n\u00de \u00bc 2E1 p 3 \u00fe 2E2q p 6 \u00fe 2E3q p 6 \u00fe 4E4q p 6 \u00fe 2E1 p 3 \u00fe 2E2q p 6 \u00fe 2E3q p 6 \u00fe E2e p 3 \u00fe E3e p 3 \u00fe 3E4q p 6 (9) For the 5-web corner element found in the MT structures of this study which is shown in Figure 11, there are also two different crushing modes associated with each half which are shown in Figure 11a,b. So the overall absorbed energy in a crushing wave in a 5-web corner element with configuration shown in Figure 11 is obtained as follows: Figure 7"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002533_j.tws.2020.107247-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002533_j.tws.2020.107247-Figure1-1.png",
        "caption": "Fig. 1. (a) Unit cell of the hollow pyramidal lattice truss core; (b) the view of the four hollow struts on a plane corresponding to the lower surface of the top face sheet.",
        "texts": [
            " However, the high temperature mechanical behaviors of metallic lattice truss core sandwich structures have been barely reported, which are potential applicability as structure having double-functionally loading and thermal feature in aerospace and nuclear industries. Therefore, the effect of elevated temperature on the mechanical properties of nickel based hollow pyramidal lattice truss structure will be investigated through theoretical analysis, experimental measurements and numerical prediction in this work. Fig. 1 presents a unit cell of the pyramidal lattice truss structure with hollow trusses. The relative density of the unit cell \u03c1 is the ratio of the lattice truss volume to that of the unit cell. For the hollow pyramidal lattice truss core, its relative density is [26]: \u03c1= \u03c0 ( d2 \u2212 d2 1 ) 2(2k + l cos \u03c9)2 sin \u03c9 (1) where the truss length l; the inclination angle between the truss and the bottom plane \u03c9, the outer diameter of the hollow truss d hollow truss wall thickness tc, the inter diameter of the hollow truss d1. The hollow truss centers can be offset by a distance of 2k. In all the designs based on CAD models of the evaluated lattice structures, the relative density \u03c1 = 0.11 was kept constant to make the results comparable. Considering the unit cell of the hollow pyramidal lattice truss core as sketched in Fig. 1, the nominal compressive modulus can be expressed as: E33(T)= sin 4 \u03c9Es(T)\u03c1 (2) A force balance along the compression axis can be used to derive an equation relating the peak strength of the hollow pyramidal lattice truss unit cell to the failure stress of the individual hollow truss. The dominant failure modes under out-of-plane compression can be considered as yielding and buckling hollow trusses, and the corresponding compressive strengths, \u03c3pk(T), can be derived as: \u03c3pk(T)= \u23a7 \u23aa\u23a8 \u23aa\u23a9 sin 2 \u03c9\u03c1\u03c3y(T), yield \u03c02sin 2 \u03c9\u03c1Es(T) ( d2 1 + d2) 4l2 , buckling (3) The critical relative density of these two failure modes can be expressed as: \u03c1critical = 2l2\u03c3y(T) ( d2 \u2212 d2 1 ) \u03c0Es ( T )( d2 1 + d2 ) (2k + l cos \u03c9)2 sin \u03c9 (4) The investigated sandwich structure with 4 \u00d7 4 hollow pyramidal lattice truss unit cells was designed using SolidWorks\u00ae 3D CAD software and manufactured using an additive manufacturing technology called Selective Laser Melting (SLM)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002600_ange.202003737-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002600_ange.202003737-Figure2-1.png",
        "caption": "Figure 2. a) The rapid deformation of MXene film when put on the palm. b) Bending downward of MXene film once a fingertip approaching. c) Flat state of MXene film when put on the palm wearing a glove. d) No locomotion for MXene film in the enclosed space filled with moisture. e) Motion process (A-F) of MXene film on a moist surface of cellulose filter paper.",
        "texts": [
            " f) Photographs of MXene film fabricated by vacuum-assisted filtration, indicating its outstanding flexibility in the inset. g) Cross-sectional SEM image of homogeneous MXene film. h) Ultrathin sections of the cross section of MXene film. i) XRD of the MXene film and MAX phase. j) Schematic of the spacing between the two neighbouring MXene nanosheets. To learn the homogeneous MXene film\u2019s response to moisture at room temperature, we first put the MXene film on our palm. Clearly, it can quickly bend in less than one second (Figure 2a), and when we place fingertip near the MXene film (non-contact) (Figure 2b), it will follow swinging with finger swing (Videos 1), demonstrating that MXene film is extremely sensitive to moisture. Nevertheless, the MXene film keeps flat state with no pronounced motion trend when put on the palm wearing a glove (Figure 2c), indicating that temperature is not the reason for the deformation. In addition, when put the MXene film in the enclosed space filled with moisture, it still maintains its initial shape and shows no locomotion (Figure 2d). The above evidences suggest that it is moisture gradients instead of 10.1002/ange.202003737 A cc ep te d M an us cr ip t Angewandte Chemie This article is protected by copyright. All rights reserved. humidity or thermal expansion that trigger the deformation of MXene film, which is consistent with the previous reports about cellulose nanofiber film[22] and GO film[8]. It is important to note that, no moisture gradient-driven motion (Figure S5, Supporting Information) is observed for both conductive graphene film and carbon nanotube film preprared using the vacuum-assisted filtration because of the lack of hydrophilic oxygen-containing surface groups. Therefore, it is no exaggeration to say that the MXene film is the only high-conductivity homogeneous film whose deformation can be driven by moisture gradients. To intuitively understand the response behavior of MXene film to moisture gradients, we randomly selected a MXene film and placed it on the surface of cellulose filter paper that can pass water vapor. This process was achieved by laying the cellulose filter paper on a 40\u00b0C water bath (Figure 2e), and then observed the deformation process of MXene film. Interestingly, the MXene film will flip automatically in a regular manner once it touches the filter paper, and whole process includes six stages (Videos 2). When the MXene film is placed on the moist filter paper, the bottom of the film will absorb the more water molecules due to the relatively high humidity of the bottom part, thereby causes the deformation resulting from asymmetric expansion and leaves the filter paper (A). Then, the barycenter of MXene film rises and reaches the critical position (B) in which the film becomes extremely unstable; thus, the film will tip over (C)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003687_a:1008389419585-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003687_a:1008389419585-Figure1-1.png",
        "caption": "Figure 1. The dynamic model of a pair of gears.",
        "texts": [
            " The parameters, \u03b7, \u03c9n, \u03b6 of the second Markov process, Equation (2), depend on the sample period 1t and are related to the rotational speed of gears. In order to study the effects of the random error of gears on dynamic behavior, the influence of the shaft stiffness is neglected and the vibration system of a pair of spur gears is considered as a stochastic system with one degree of freedom. Furthermore, we consider the errors as consisting of two components: harmonic and random. So, the equation of motion of the dynamic gear pair model shown in Figure 1 is given as Mx\u0308 + C(t)x\u0307 +K(t, x)x = W + Fr + Fh, (14) where x denotes the relative displacement between the driving and the driven gears along a line of action, see Figure 1a. M denotes the equivalent mass of a gear transmission system and is calculated by M = M1M2/(M1 +M2), where M1, M2 are, respectively, the equivalent masses of the driving and the driven gears. C is the damping coefficient and is calculated by C(t) = 2\u03b60 \u221a MK(t, x). \u03b60 is the damping ratio and assumed to be 0.07. K(t, x) is the time-variant gear mesh stiffness along a line of action of two gears. W is the static force between two gears along a line of action. Fh is the force excited by harmonic errors eh, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002961_tmag.2021.3082686-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002961_tmag.2021.3082686-Figure1-1.png",
        "caption": "Fig. 1. Structural design for ultra-high speed vacuum pipeline transportation. Lift and guidance forces are produced by the interaction of the permanent magnet Halbach array with aluminum plate; permanent magnet linear synchronous motor drives the maglev train. Subgraph (a) shows the overall structure. The maglev train is \u2018wrapped\u2019 in the reinforced concrete bed; linear motor stator is fixed on T-shaped concrete bed; four aluminum plates are embedded symmetrically on both sides of the track. Subgraph (b) shows the structure of the maglev train. Eight groups of Halbach permanent magnet arrays are symmetrically installed on both sides and at the bottom of the suspension frame, facing the aluminum plates.",
        "texts": [
            " Secondly, a brief mapping is established between the electromagnetic forces and structural parameters by omitting high order harmonic components, laying the foundation for the structural optimization. The optimal wavelength-to-gap and thickness-to-gap ratio are proposed by solving first order partial differential equations. Then, the finite element simulation models are established to verify the correctness of the optimal structural parameters. Finally, a rotary experimental device is built to validate the validation of the analytical electromagnetic forces. II. STRUCTURAL DESIGN OF THE ELECTRODYNAMIC SUSPENSION SYSTEM The structure of the electrodynamic suspension system is shown in Fig. 1. The maglev train mainly include the train body and suspension frame. The suspension frame is made of high strength aluminum alloy to support the car body. Eight groups of permanent magnet Halbach arrays are symmetrically Authorized licensed use limited to: Tencent. Downloaded on June 30,2021 at 09:16:13 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2021 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_14_0001883_j.oceaneng.2020.107826-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001883_j.oceaneng.2020.107826-Figure1-1.png",
        "caption": "Fig. 1. DTC hull (yellow) with Subdivision Surface control polygon (black) and FFD control volume (blue). (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " The control polygon of the Subdivision Surface representing the hull is made of 228 control points. For the purposes of the following analysis we have focused on the bulb geometry, that is described by a control polygon made of 27 control points. In addition an orthogonal FFD control volume made of 3 planes for each direction, hence corresponding to 27 independent control points, is applied to the bulb surface. Both the Subdivision Surface control polygon and the FFD control volume are shown in the perspective view of the DTC hull in Fig. 1. In this particular case the control points of the Subdivision Surface and that of the FFD transformation are the same number. The geometric relation created between the Subdivision Surface and the FFD transformation allows to use simpler FFD boxes e.g. with respect to a conventional FFD approach, by keeping the same transformation complexity. In this particular case for example, by using a conventional FFD approach, i.e. applying the transformation directly to the vertices of the surface mesh, in order to ensure \ud835\udc362 continuity along a specific direction two layers of control points of the FFD box would have to be restrained from movements in that direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003812_(sici)1521-4109(199906)11:7<505::aid-elan505>3.0.co;2-j-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003812_(sici)1521-4109(199906)11:7<505::aid-elan505>3.0.co;2-j-Figure4-1.png",
        "caption": "Fig. 4. CV of 2.3 mM H2Q in 0.1 M phosphate buffer pH 7.4 on bare electrode (curve 1) and on SME (curve 2). v 50 mVys.",
        "texts": [
            " At lower scan rates the linear relationship versus v1y2 (panel A) indicates a diffusive control of the electrochemical process; in these ranges of v the peak intensity changes linearly with FCA concentration (curve not shown). At higher values such relationship is lost, and iox is now a function of v as shown by panel B where log iox versus log v is reported (slope 0.97). Such relationship indicates an eT controlled by a surface process [6] like adsorption, complexation etc. We tested also the behavior of SME towards HQ that, like FCA, is perfectly permeable through the bCD cavity, but forms inclusion complexes with bCDs weaker [14] than FCA. In Figure 4 the CV of HQ at bare electrode and at SME are shown as an example. Although some reversibility is lost, actually the signal on SME, contrarily to bare electrode, is much more stable: CV does not change with time as for bare Au electrode (due to variable interfacial arrangement). This is probably due to the monolayer at the surface which inhibits the spontaneous oxidation processes on the electrode surface (consequently not passivated [15 \u00b118]). For HQ the iox changes linearly both with the concentration and v1y2 (curves not shown), as expected for a process controlled by diffusion",
            " 7 Preparation and Electrochemical Characterization of a b-Cyclodextrin-Based Chemically Modi\u00aeed Electrode 507 higher concentration or longer cycling produced passivated surfaces while the contrary was not suf\u00aecient to cover all pinholes. The ef\u00aeciency of HIAA treatment was tested by species as HQ and FCA which can permeate the bCD cavity and by Fe(CN)\u00ff3 6 that for steric reasons is unable to permeate into the bCD cavity [21]. In particular, an optimal treatment does not produce any substantial change on the CV of FCA, while it determines a partial recovery of reversibility for the HQ: the CV becomes very similar to that on bare electrode (curve 1 of Fig. 4). On the contrary, with Fe(CN)\u00ff3 6 the CVs (curve not shown) are similar to the blank (no signal), as a result of the surface preventing the eT. These results indicate that HIAA treatment, under suitable controlled conditions, allows one to cover only pinholes meanwhile maintaining free the bCD cavities. In addition the pinhole coverage was evidenced also by a slight decrease in the double layer capacitance. We tried using the SME as a channel sensor [22\u00b124] to detect the presence in the working solution of nonelectroactive substances (but forming highly stable inclusion complexes with bCD)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000489_2016-01-1560-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000489_2016-01-1560-Figure2-1.png",
        "caption": "Figure 2. (a) Orientation of contact normal determined by angles \u03b1 and \u03b2, and (b) double-inclined plane representation of contact surface showing coordinate systems CS1 and CS2; (c) global view of ball in BNA showing global coordinate system and coordinate systems CS2 and CS3.",
        "texts": [
            " To determine the elastic deformations of the screw and nut, the contact force Pi must be transformed from a local coordinate system defined along the contact normal of each ball i to a global coordinate system defined for the BNA [14,15,16]. The procedure for determining this local-to-global coordinate system transformation is exactly the same for the screw and nut. Therefore, for the sake of brevity, only the transformations for the screw are highlighted in this paper. Details on the transformations for the nut can be found in [14,15,16]. Figure 2(a) shows the x, y, z global coordinate system for the BNA, as well as the local coordinate system (CS1 \u225c x1, y1, z1) whose z-axis lies along the contact normal for ball i. Angle \u03b1 shown in the figure is the lead angle of the screw, given as (2) where l is the lead and d is the pitch diameter of the screw; angle \u03b2 is the contact angle at the ball contact point. Using these two angles, the ball-screw contact surface can be represented as a double-inclined plane as depicted in Figure 2(b). An intermediary coordinate system (CS2 \u225c x2, y2, z2) is defined such that its z-axis points along the axis of the screw and its y-axis lies along the radius of the screw. The transformation between z1 and CS2 (i.e., T2-z1) is obtained using current frame rotation operators about the y and x axes, and then multiplying them by a unit vector in the z1 direction, as (3) where rot represents a current-frame rotational operator about the axis specified by its subscript [14]. Figure 2(c) shows the global view of a ball in a BNA. Another intermediary coordinate system, CS3 \u225c x3, y3, z3, is defined with its origin at the contact point and its x, y and z axes are parallel with those of the global coordinate system. The azimuth angle \u03d5 determines the position of the ith ball on the circumference of the screw. Accordingly, the transformation T3-2(i) from CS3 to CS2 simply involves a rotation by the amount \u03d5i+\u03c0/2 about the z axis; i.e., (4) The transformation between the contact normal direction z1 and CS3 is therefore obtained as follows (5) Using transformation matrix T3-z1(i), the contact force Pi is transformed to an equivalent force F3S(i) in the intermediary coordinate system CS3 as (6) To describe the displacement of the screw under the contact forces, the screw is modeled as a Timoshenko beam",
            " The proposed low-order (beam finite element) model which considers axial, torsional and lateral deformations of the screw and nut iii. A high-order (3D finite element) model which considers axial, torsional and lateral deformations of the screw and nut, created using Solidworks finite element package Figure 4 compares the load distribution of the balls using the three models under investigation. Note that, because preload is achieved using oversized balls, each ball contacts the raceway of the screw at two points, i.e., at a lower (L) and upper (U) contact point [14] (as opposed to the single lower contact point shown in Figure 2). As seen from the figure, the load distributions predicted by the Solidworks model for the L and U contact points are in good agreement with those of our proposed model; both models show that the contact load distributions of the balls are not uniform for the L and U contact points. However, Mei et al.\u2019s model predicts uniform load distributions on all balls, which is inaccurate. The weakness of Mei et al.\u2019s model lies in the fact that it cannot capture the coupling between axial, torsional and lateral deformations [14,15,16] created by the oversized balls"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002352_j.mechmachtheory.2020.103945-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002352_j.mechmachtheory.2020.103945-Figure4-1.png",
        "caption": "Fig. 4. \u201c(4R)\u201d modules: (a) \u201c \u02dc R u 1 / \u0303 R u 2 \u201d, (b) \u201c \u02dc R u 1  \u0303 R u 2 \u201d, (c) \u201c \u2194 R u 1 \u22a5 \u0303 R u 2 \u201d, and (d) \u201c \u02dc R u 1 \u22a5 \u2194 R u 2 \u201d.",
        "texts": [],
        "surrounding_texts": [
            "W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx 3\nwhere M sl and n sl denote respectively the DOF of the external single-loop linkage and the number of R pairs of the linkage, and C cc and n cc denote respectively the number of the constraints contributed by the coupling chain and the number of R pairs of the chain. From Eq. (1) , n sl and n cc can be enumerated in detail, listed in Table 1 .\nIn the recent work [34 , 35] , two classes of deployable units with n sl = 8 and n cc = 5, and n sl = 9 and n cc = 4 have been synthesized. The remaining three classes of units with n sl = 10 and n cc = 3, n sl = 11 and n cc = 2, and n sl = 12 and n cc = 1 will be developed here.\nA two-layer and two-loop deployable linkage unit needs to have a plane-symmetrical structure, so its R pairs need to be arranged symmetrically with respect to the symmetrical plane (SP). As a result, for the deployable unit with n sl = 10 and n cc = 3, there exist only two kinds of structure types, denoted as \u201c{4R} [3R] (6R)\u201d and \u201c{6R} [3R] (4R)\u201d, shown in Fig. 1 (b) and (c), in which FB denotes the fixed base, and MP1 and MP2 denote the two links connecting the coupling chain with the external loop linkage. A \u201c{4R} [3R] (6R)\u201d unit can be decomposed into a 4R lower module (denoted as \u201c{4R}\u201d), a 3R coupling chain module (denoted as \u201c[3R]\u201d), and a 6R upper module (denoted as \u201c(6R)\u201d). Similarly, a \u201c{6R} [3R] (4R)\u201d\nunit can be decomposed into a lower module \u201c{6R}\u201d, a coupling chain module \u201c[3R]\u201d and an upper module \u201c(4R)\u201d.\nFor the deployable unit with n sl = 11 and n cc = 2, there exists only one structure type shown in Fig. 1 (d), denoted as\n\u201c{6R} [2R] (5R)\u201d. Such a unit can be decomposed into three modules \u201c{6R}\u201d, \u201c[2R]\u201d and \u201c(5R)\u201d.\nFor the deployable unit with n sl = 12 and n cc = 1, there is only one structure type shown in Fig. 1 (e), denoted as \u201c{6R} [1R]\n(6R)\u201d. Such a unit can be decomposed into three modules \u201c{6R}\u201d, \u201c[1R]\u201d and \u201c(6R)\u201d.\n2.2. Structure modules of new deployable linkage units\nThe above linkage units involve two kinds of lower modules\u201c{4R}\u201d and \u201c{6R}\u201d, three kinds of coupling chain modules\u201c[3R]\u201d, \u201c[2R]\u201d and \u201c[1R]\u201d, and three kinds of upper modules\u201c(6R)\u201d,\u201c(5R)\u201d and \u201c(4R)\u201d. Synthesis of those modules is the important basis of developing the three classes of linkage units. For convenience, in what follows, \u201c \u2194\nR i \u201d denotes that R i is\nlaid along the centerline of a link connected by the pair, \u201c \u02dc R i \u201d denotes that R i is perpendicular to the centerline of a link connected by the pair. \u201cR i \u201d denotes that R i is in the SP. In addition, notations \u201c/\u201d, \u201c\u22a5 \u201d and \u201c \u201ddenote the parallel, perpendicular and noncoplanar relations between joint axes, respectively.\nBy considering plane-symmetry and enumerating different joint layouts, topological structures of each kind of module can be synthesized, and the related results are presented as follows, in each of which the layouts of only R pairs in the left\nof the SP and in the SP.\n2.2.1. Coupling chain modules\nThere are three \u201c[3R]\u201d modules denoted as \u201c \u02dc R c1 / \u0303 R c2 \u201d, \u201c \u02dc R c1  \u0303 R c2 \u201d and \u201c \u2194 R c1 \u22a5 \u0303 R c2 \u201d, shown in Fig. 2 (a\u2013c). Besides, there are only one \u201c[2R]\u201d module denoted as \u201c \u02dc R c1 \u201d and only one \u201c[1R]\u201d module denoted as \u201c \u02dc R c1 \u201d, shown in Fig. 2 (d) and (e), respectively.\n2.2.2. Lower modules\nThere are two kinds of \u201c{4R}\u201d modules, \u201c \u02dc R l1  \u0303 R l2 \u201d and \u201c \u02dc R l1 \u22a5\n\u2194 R l2 \u201d, shown in Fig. 3 (a) and (b), and there are three kinds of\n\u201c{6R}\u201d modules, \u201c \u02dc R l1 \u22a5\n\u2194 R l2 \u22a5 \u0303 R l3 \u201d, \u201c \u02dc R l1  \u0303 R l2 \u22a5 \u2194 R l3 \u201d and \u201c \u02dc R l1  \u0303 R l2 / \u0303 R l3 \u201d, shown in Fig. 3 (c\u2013e), respectively.\nPlease cite this article as: W.-a. Cao, Z. Jing and H. Ding, A general method for kinematics analysis of two-layer and twoloop deployable linkages with coupling chains, Mechanism and Machine Theory, https://doi.org/10.1016/j.mechmachtheory. 2020.103945",
            "4 W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx\nRu1 MP1 MP2\nRu2 Ru4\nRu3 Ru1 MP1 MP2\nRu2 Ru4\nRu3 Ru1\nMP1 MP2\nRu2 Ru4\nRu3\nRu2\nRu1 Ru3\nRu4\nMP1 MP2\n(a) (b) (c) (d)",
            "W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx 5\nFurther, some units should be excluded due to their structural defects, such as local DOFs and undesired angle ranges[35]. Finally, there are fifteen available \u201c{4R} [3R] (6R)\u201d deployable units, listed in Table 2 , in which two typical units are shown in Fig. 7 (a) and (b).\nSimilarly, seventeen available \u201c{6R} [3R] (4R)\u201ddeployable units can be obtained, listed in Table 3 , in which two typical\nunits are shown in Fig. 7 (c) and (d).\nBased on the combination of three kinds of \u201c{6R}\u201d modules, one kind of \u201c[2R]\u201d module and three kinds of \u201c(5R)\u201d modules, there are nine \u201c{6R} [3R] (5R)\u201d units in total. After excluding those units with structural defects, six\navailable\u201c{6R} [2R] (5R)\u201d units can be obtained, listed in Table 4 , in which two typical units are shown in Fig. 8 (a) and (b). Similarly, there are also six available \u201c{6R} [1R] (6R)\u201ddeployable units, listed in Table 5 , in which two typical units are shown in Fig. 8 (c) and (d).\nPlease cite this article as: W.-a. Cao, Z. Jing and H. Ding, A general method for kinematics analysis of two-layer and twoloop deployable linkages with coupling chains, Mechanism and Machine Theory, https://doi.org/10.1016/j.mechmachtheory. 2020.103945"
        ]
    },
    {
        "image_filename": "designv11_14_0003672_ieeestd.2000.91147-Figure61-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003672_ieeestd.2000.91147-Figure61-1.png",
        "caption": "Figure 61\u2014Influence of load on mho relay characteristic",
        "texts": [
            " A large value of fault resistance causes the mho relay at terminal G to sense a fault that should be within the reach of the relay as being outside of the reach. If the fault is outside the reach in the forward direction, the relay may misoperate if the argument is negative. The impact of the combined effect of load and fault resistance on operation of the distance relays depends on the relay characteristic. An example of a typical quadrature-polarized ground distance mho unit has been analyzed for a phase \u201ca\u201d to ground fault. The operating characteristics for load and no-load cases are shown in Figure 61. Input signals for a phase comparator unit that will operate when the operating signal lags the polarization signal by 90\u2013270\u00b0 are where VGfa is the \u201ca\u201d phase voltage at Bus G during the fault; Isfa is the \u201ca\u201d phase current from Bus G into the fault; ZL0 is the zero-sequence line impedance between Buses G and H; ZG V Gf Isf -------- mZL Isf\u00d7 R f I f\u00d7+ Isf ---------------------------------------------- mZL R f I f I sf -----    mZL=\u00d7 R f ks\u00d7+ += = = V Gfa Isfa ZL0 ZL1\u2013 ZL1 ---------------------- Isf 0\u00d7+ ZC\u00d7 operating signal\u2013 j V Gbc\u00d7 polarization signal Authorized licensed use limited to: ULAKBIM UASL - Anadolu Universitesi",
            " Downloaded on May 01,2014 at 23:43:40 UTC from IEEE Xplore. Restrictions apply. IEEE Std C37.113-1999 IEEE GUIDE FOR PROTECTIVE RELAY 80 Copyright \u00a9 2000 IEEE. All rights reserved. Z L1 is the positive sequence line impedance between Buses G and H; I sf0 is the zero-sequence current from Bus G into the fault; Z C is the relay setting impedance; j is the square root of negative 1; V Gbc is the voltage at Bus G between phase \u201cb\u201d and phase \u201cc.\u201d The relay operating characteristic has an offset, represented by length AL in Figure 61. The offset impedance, Z os , may be expressed as follows: (5) where Z s0 and Z s2 are the zero and negative-sequence source impedances; Z L0 is the zero-sequence line impedance between Buses G and H; Z L1 is the positive sequence line impedance between Buses G and H; d s2 and d s0 are the current distribution factors in the negative and zero-sequence networks, respectively. They represent the ratio of the change in the current at the terminal caused by the fault (fault current minus prefault load current) to the total fault current; I L is the prefault load current. The offset increases with larger value source impedances, which enables adjustment of the characteristic to larger fault resistance. The offset also depends on the load current, I L . In Figure 61, the no-load case ( I L = 0) Zos 2ds2 Zs2\u00d7 Ds0 Zs0\u00d7+ 2ds2 ZL0 ZL1 -------- ds0 IL Isf 0 --------+\u00d7+ ------------------------------------------------------= Authorized licensed use limited to: ULAKBIM UASL - Anadolu Universitesi. Downloaded on May 01,2014 at 23:43:40 UTC from IEEE Xplore. Restrictions apply. IEEE APPLICATIONS TO TRANSMISSION LINES Std C37.113-1999 Copyright \u00a9 2000 IEEE. All rights reserved. 81 is compared to the load case. The offset for the no-load case is represented by the distance from point A to point K"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001071_s0263574719000924-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001071_s0263574719000924-Figure4-1.png",
        "caption": "Fig. 4. Simulation model of free-floating space robot system.",
        "texts": [
            " In this way, the two processes need only one calculation during one visual sampling period, which means that the two processes would only have influence on one or two control periods. Besides, when the calculated processes are programmed by C-code, it is verified that the processes can both be finished within 2 ms. In order to eliminate the influence on the further control caused by the two processes, the two parts are calculated in two separate control periods in this paper, and thus, the time requirements are satisfied. In other words, the method can be applied to the practical situations. The dual-arm FFSR system for the simulation is shown in Fig. 4(a), and it is mainly composed of two 7-DOF space manipulators and a satellite base. The D\u2013H coordinate system is shown in Fig. 4(b) and corresponding D\u2013H parameters are shown in Table I. In the D\u2013H coordinate system, the base coordinate, Ob, does not align with the manipulator installation coordinate, O0, and the transform https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574719000924 Downloaded from https://www.cambridge.org/core. Nottingham Trent University, on 15 Jul 2019 at 07:15:28, subject to the Cambridge Core terms of use, available at rule is also presented in Table I. In Table II, the inertia parameters are listed, which are indispensable for the general Jacobian and dynamics"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure3-1.png",
        "caption": "Fig. 3. The geometric model for CCHGP units. (a) 2D control area separation, for a particular angular position \u03d5, done by the numerical geometrical model; (b) 3D Tooth-space geometry and the sealing surfaces in CCHGP meshing zone. The sealing surfaces are created by the meshing of transverse rotor profiles separate the tooth-space into upper and lower control volumes (CVs). The red arrows illustrate representative flow passages between meshing volumes. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " Among different modules employed, the fluid/pressure dynamics module and its associated part of the geometric module have been discussed in detail in the authors\u2019 previous work focused on kinematics and fluid dynamics [20]. In the following part of this section, some core concepts of the fluid dynamics model will be recalled. The lumped-element fluid dynamics model formulation takes advantage of the fact that in CCHGP type units, the fully conjugate gear profile realizes a separation of the tooth space volumes in the axial direction z (Fig. 3). This transverse section where the tooth-tip of one gear reaches the bottom of the tooth-space of the other gear provides an axial \u2018sealing surface\u2019, which cuts the tooth-space volume into two control volumes, or CVs (see Fig. 3, the control volume whose centroid is located at a higher z-position is called upper CV, and the other CV is called lower CV). Such \u2018sealing surface\u2019 moves up or down (along the z-axis) as the gear set rotates, depending on the direction of the helix. Therefore, this sealing surface can be considered a moving solid boundary that contributes to the displacing action of the CCHGP. The fluid exchange between these two sub-volumes, in the axial direction, is limited by this sealing surface and high pressure-difference can exist between the control volume above the sealing surface (i",
            " lumped-element model, has been widely applied to the modeling of a large variety of hydraulic machines in the fluid power community [21,29\u201331]. The pressure build-up for each CV can be derived from mass conservation, which leads to the equation: X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 dp dt = K Vcv ( 1 \u03c1 \u2211 m\u0307 \u2212 dVcv dt ) (1) where bulk modulus K = \u03c1 dp d\u03c1, and the mass fluxes m\u0307 are evaluated for each hydrodynamic connection between control volumes, depending on the physics represented by that connection. Some typical flow passages are highlighted by red arrows in Fig. 3b. For CCHGP units, all the connections except the leakages through the lateral lubrication gap are modeled with the orifice equation with Reynolds number modification [29], which is written as: m\u0307orif = \u03c1cqA \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 2|\u0394p| \u03c1 \u221a (2) where \u03c1 is the averaged density evaluated at the averaged pressure between inlet and outlet chamber for each hydrodynamic connection, A is the opening area, and cq is the discharge coefficient. The calculation of geometrical opening area A between a pair of CVs at particular angular positions through a geometric module is explained in [20] in detail",
            " From the fluid-pressure dynamics model, the pressure values are solved for each of these control volumes. The determination of force and moments of fluid-pressure forces acting on the radial surface of gears also requires geometrical information about the application areas and moments of each control area on the gears. The application areas and the coordinates of application points are X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 obtained by integrating the 2D geometric information along the axial direction of the gear. Fig. 3a indicates that each control area is in contact with at least one gear profile, and the region of contact varies with gear rotation angle. In particular, the Gear 1 TS area is in contact with Gear 1 all the time, while only in contact with Gear 2 when it has deformed morphology; the same is true for Gear 2 TS area: it is in contact with Gear 2 all the time while only in contact with Gear 1 when it is deformed. The differential force and moment can be obtained from the 2D geometric model by tracking the position of the interfacial curve representing the interface between control volume and 2D transverse gear profiles",
            " For helical gear pumps, the profile angle changes linearly with axial position: dz = H \u0398 d\u03d5 (20) Substituting Eq. (20) into Eq. (18), all geometric groups can be integrated. In the Eqs. (14), (15) and (16), the axial positions used as integration boundaries (namely, z1, z2) are pre-determined from the meshing geometry between two gears. The integration boundaries for different control volumes are shown in Table 2. For inlet/outlet variable volume Vin and Vout, the integration is from the bottom to the top of the gear pump, whereas for TS volumes (upper and lower CVs for two gears, as shown in Fig. 3b), the integration is separated by the position of the sealing surface. zG1 and zG2 are the axial positions of the sealing surface for Gear 1 and Gear 2, respectively. X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 The fluid-pressure distribution within the lateral gap between the gear axial surface and the lateral bushing (i.e. bearing blocks) are solved with the Finite Volume Method (FVM) with an unstructured solver modified based on the open-sourced package OpenFOAM",
            " The force and moments that the lubrication film exerts on the gear axial surface can be integrated as follows (subscript 2 stands for the fluid-pressure force on the axial gear surface): Fz,p2 = \u03c3 \u222b A p dA Mx,p2 = \u03c3 \u222b A p y dA My,p2 = \u2212 \u03c3 \u222b A p x dA (26) while the friction force (in the x-y plane) and torque loss by shearing is given by X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 Fxy,p2 = [ Fx,p2 Fy,p2 ] = \u222b A \u03c4 dA Mz,p2 = \u222b A |r \u00d7 \u03c4| dA (27) where \u03c4 = \u2212 h 2 \u2207p+ \u03bc h v r = [ x y ] (28) The solution of Eq. (21) depends on the pressure boundary condition. For each face on the boundary, the geometric model (Fig. 3) determines which control area it belongs to, and the pressure values are taken from the fluid-pressure dynamics model described in Section 3. Four gaps, namely, the top and bottom gaps for Gear 1 and Gear 2, are coupled into the system and solved for each time step, and the calculated force and moments are coupled back to the gear position solver. Two types of contact forces are discussed in this section, denoted as Fc1 and Fc2. Fc1 is the driving contact force that occurs on the driving flank of the gear teeth, along the line of action, to overcome the load generated by the fluid pressure",
            " A stronger vibration can be observed in the angle \u0394\u03b8y, which results in a fluctuation in the center distance between two gears (Fig. 21b). At the highpressure high-speed condition operating condition, the transverse force vectors on two gears are pointing outwards (Fig. 17), therefore the center distance is consistently higher than the pitch diameter 2rp. The fluid dynamics indication of such condition (i > 2rp) is discussed in detail in [20]. Simply speaking, it introduces a gap to the axial sealing surface (see Fig. 3), which increases the outlet flow ripple, but attenuates the internal pressure peak (Fig. 17a). A larger tilt is found for Gear 1 than for Gear 2, resulting in a smaller center distance on the top side than on the bottom side (i.e. i1 > i2, see Fig. 21b). This is because the moment My on Gear 1 is greater than that on Gear 2, as discussed earlier (Fig. 20a). However, this pattern changes in the high-pressure low-speed condition. With the low-speed rotation of the rotor shaft, the lubrication films in the journal bearings have a weaker bearing function, resulting in a different shaft position and a different circumferential TS-pressure distribution around gears"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000719_mma.4210-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000719_mma.4210-Figure3-1.png",
        "caption": "Figure 3. Trajectory of the end effector for a flexible sliding rod.",
        "texts": [
            " Equations of motion The Lagrange formulation [19] to derive equations of motion can be expressed as d dt @T @Psi C @T @Psi D Ri , i D 1,2 where Ri are the generalized forces, s1 D , s2 D rare the time dependent generalized coordinates and the subscript i D 1,2 represents the number of the generalized forces/coordinates. The total kinetic energy of the system can be expressed as T D NX iD1 Ti D 1 2 NX iD1 miv 2 Ci C ICi! 2 i where Ti is the kinetic energy, ICi is the mass moment of inertia, and !iis the angular velocity of each rod. The displacement of the end effector along the motion path, when the sliding rod SP is flexible (Figure 3), can be calculated as in [20]. Then, the equations of motion (based on the mode shapes of a clamped-free beam) can be expressed by \u0152M 8< : Rs Ru R 9= ;C \u0152H 8< : Ps Pu P 9= ;C \u0152G 8< : s u 9= ; D 8< : 0 F T 9= ; Copyright \u00a9 2016 John Wiley & Sons, Ltd. Math. Meth. Appl. Sci. 2016 where Fis the axial force that slide the flexible road in the rigid guide OE, T is the torque that rotate the guide, and u is the un-deformed part of the sliding rod (Figure 3). The matrices \u0152M \u0152H and \u0152G are defined as in [20] by \u0152M D f .l0, , A, Me, u, \u0152I , \u0152Ci , \u0152Ci / , \u0152H D f .l0, , A, Me, u, Pu, \u0152I , \u0152Ci , \u0152Ci / , \u0152G D f .l0, , A, E, Me, u, Pu, Ru, \u0152I , \u0152Ci , \u0152Ci / , where \u0152Ci denote a square matrix [20]. Two numerical examples to generate the piecewise polynomial trajectory of the end effector as well as the trajectory of the non-active link end, which constrains the motion are considered. The possible locations of the base as well as the phase diagram (position vs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000082_icuas.2019.8798198-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000082_icuas.2019.8798198-Figure1-1.png",
        "caption": "Fig. 1. Inertial and body frame; angles; force and torques",
        "texts": [
            " Section III deals with the design of the B controller based on low and high frequency sensitivity bounds. These same bounds are set as objectives in the design of \u03bc-synthesis control in section IV. Section V presents PID design which aims to reach the same rise time as B control. Section VI compares simulation results. Some future directions of research will be proposed in the conclusion. Non-linear quadcopter modelization is already largely investigated in literature as in [12-14]. It is generally represented in the body frame attached to the quadcopter as in Fig. 1. The second frame , called inertial one, indicate the starting point of the quadcopter. With respect to the inertial frame linear accelerations on x, y, z can be written as follow: (1) Where C(.) and S(.) are the respective abbreviations of cos(.), and sin(.). Angles , , around x, y, z axis are shown in Fig..1. Air friction on the chassis along x, y, z axis is represented by coefficients Axx, Ayy, Azz; m is the mass of the drone and g the gravity constant. The force Tz is generated by the four propellers systems. Angular accelerations around , are given by: ! \" ! (2) Where L, M, N represent torques around x, y, z as shown in Fig. 1. Inertial coefficients along x, y, z axis are respectively Ixx, Iyy and Izz. D B. Linear approximation model By considering steady flight we apply Taylor linearization on (1) and extract its nominal part (assuming and ), which leads to the following equation: 978-1-7281-0332-7/19/$31.00 \u00a92019 IEEE 344 Where # is the force generated by propellers in steady flight. Control structure in Fig. 2 illustrates two nested feedback loops: an inner loop and an outer loop. The inner loop with the faster bandwidth is dedicated to the control of orientation dynamics on [ ]T through L, M, N variables; desired input d is set by the user whereas d and d are given by x and y translation controllers from the outer loop in charge of the control of the vehicle position along a given trajectory"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001956_j.ijmecsci.2020.106137-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001956_j.ijmecsci.2020.106137-Figure13-1.png",
        "caption": "Fig. 13. Influence of strut waviness distribution on ultimate tensile strength and ductility of triangular lattices. Effect of A \u221e/ \ud835\udcc1, A m / \ud835\udcc1 and \ud835\udf06/ \ud835\udcc1 on normalised UTS ?\u0302? for triangular lattices with notch sizes 13 (a) \ud835\udc5b b = 1 and 13 (b) \ud835\udc5b b = 4 , and on ductilities \ud835\udf00 \u221ef of lattices with 13 (c) \ud835\udc5b b = 1 and 13 (d) \ud835\udc5b b = 4 .",
        "texts": [
            " Here, the struts adjacent to the notch fail first, while the wavy struts next to them are not stretched out completely and are far from reaching the tensile necking strain. .3.2. Triangular lattices A similar study has been performed on triangular lattices, with a repesentative square periodic RVE of size 2 D \u00d7 2 D ( \ud835\udc37\u2215 \ud835\udcc1 = 12 ) and notches f length \ud835\udc4e \u2215 \ud835\udcc1 = 0 . 6 and 1.7, corresponding to \ud835\udc5b b = 1 and 4, respectively Fig. 7 ). The graded waviness of struts is again defined by Eq. 4 . The esign maps of Fig. 13 show normalized macroscopic tensile strength nd ductility for triangular lattices over a wide range of waviness distriution A \u221e/ \ud835\udcc1, A m / \ud835\udcc1 \u2208 [0, 0.06], \ud835\udf06/ \ud835\udcc1 \u2208 [0, 10], for the cases of a notch s characterized by \ud835\udc5b = 1 and \ud835\udc5b = 4 . b b The imposition of graded waviness onto the triangular lattice derades the tensile strength for \ud835\udc5b b = 1 and only leads to a 1% increase in ensile strength for \ud835\udc5b b = 4 , see Figs. 13 (a) and 13 (b). This is in contrast o the significant elevation in tensile strength for the hexagonal lattice, ecall Figs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001759_j.mechmachtheory.2020.103995-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001759_j.mechmachtheory.2020.103995-Figure10-1.png",
        "caption": "Fig. 10. Trajectories of points on the linkage. (a) is front view; (b) is left view; (c) is top view.",
        "texts": [
            " For the kirigami form corresponding to the double crank, the angle \u03d5 increases all the time, but it will go through the bifurcation point where the panels p 1 and p 6 were coplanar with \u03d5= 180 \u00b0. This can be interpreted as kirigami form of type I can be directly bifurcated into type II through the bifurcation points B 3 and B\u2019 3 in Fig. 7 . The characteristics of the linkage are inherited by its kirigami form. Here, of particular interests are crank-rocker linkage and its kirigami form. The trajectories on the linkage are monitored as shown in Fig. 10 , while the trajectories on the kirigami form are plotted in Fig. 11 during the whole cycle of motion when panel p 1 is fixed. The subscript i of M i or N i indicates that the point is on the revolute joint i or the crease z i . Points located on the creases i in the kirigami form have similar trajectories to the points on the revolute joint i in the linkage. For example, M 5 and N 5 are different points on the revolute joint 5 of the linkage and crease z 5 of the kirigami form, respectively. Both the trajectories of M 5 and N 5 are circular arcs despite different starting point and arc length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000204_j.autcon.2019.102996-Figure17-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000204_j.autcon.2019.102996-Figure17-1.png",
        "caption": "Fig. 17. Subpart A: (a) outlining curve of the notch taken from the geometry of the hub; (b) notch geometry in the correct direction; (c) extrusion of notch geometry and completion of subpart A.",
        "texts": [
            " The same procedure applies, automatically, to any other hub with a different number of notches. After placing the hub in each mesh node, we can generate the adjusted geometry of the interconnecting parts. The interconnecting part geometry subdivides into three subparts: subpart A of the geometry depends on notch location; subpart C depends on element location, and subpart B connects A to C. Fig. 16 shows these subparts. 3.3.8.1. Generation of subpart A. Subpart A depends on the geometry and positioning of the hub notch. Fig. 17a shows the geometry of the hub notch with its outline and its direction vector. The outline of the notch may be modified according to the profile available. Next, the curve is extruded to form the surface of subpart A (Fig. 17b). The hub axis vector guides the extrusion of the subparts related to the corresponding hub. This step completes after top and bottom surfaces cap the notch geometry (Fig. 17c). 3.3.8.2. Generation of subpart B. Subpart B joins a rectangular section from subpart A and a circular section from subpart C. We simplified the circular and rectangular sections into segments of curves to avoid geometry errors in the program. Fig. 18a shows this simplification and subsequent surface generation. All quarters join into one surface; then, the joined surface connects to subparts A and C (Fig. 18b). 3.3.8.3. Generation of subpart C. The segments of mesh lines depicted in Fig. 12a are base for generation of subparts C, which are pipes with predefined radiuses"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003679_s0167-8922(08)70490-5-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003679_s0167-8922(08)70490-5-Figure9-1.png",
        "caption": "Fig. 9 Shell surface temperature. Diesel engine.",
        "texts": [
            " The maximum oil film temperature is equal to the maximum shell temperature almost throughout the whole cycle. Only at the pressure peaks the oil film temperature rises to a maximum of 126.5 deg-C. The minimum oil temperature is 657 always the temperature of the oil in the oil groove (97-deg-C). This results in a maximum difference of oil temperatures in the bearing of 29.5 deg-C (Fig. 7, Fig. 8). The maximum temperature rise from the inlet temperature of 90 deg-C is 36.5 deg-C. On the shell surface there is a temperature variation of 28 deg-C, with the minimum again a t the oil groove, Fig. 9. 3.2. Main bearing of a gasoline engine The Finite Element structure of this bearing is depicted in Fig. 1. Peak oil film pressure. (Fig. 10) Comparing the peak oil film pressure for pressure and temperature dependent viscosity and density ( q = q(p,T)) with the pressure for constant viscosity and density ( q = ~ lp=lbar ,T=120deg-c ) one can see that the absolute maximum over the engine cycle hardly changes for the two variants. Nevertheless, significant differences in the peak oil film pressure a t 700-60, 165- 180, 310-385 and 500-550 deg-CA can be observed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002721_j.aej.2021.01.012-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002721_j.aej.2021.01.012-Figure1-1.png",
        "caption": "Fig. 1 Worm wheel machining system.",
        "texts": [
            " The complete mathematical models of worm gear sets enable the performing of gear meshing analyses, including tooth contact analysis (TCA). Additionally a mathematical description of surfaces of worm gear sets provide generating computer-aided design (CAD) models, which can be used for e.g. finite element analysis. 2. Machining system of the worm wheel The worm wheel is machined with a cutting worm hob. The mathematical model of the tool has a continuous surface limited by the cutting edges. There are no intermediate cutting edges. The tool and worm wheel blank axes are fixed at an angle of 90 (Fig. 1). The static coordinate systems of the tool S1 x1y1z1\u00f0 \u00de and worm wheel blank S2 x2y2z2\u00f0 \u00de are introduced. These can be treated as systems connected with the body of the machine tool. Movable systems are introduced: x 0 1y 0 1z 0 1 of the tool and x 0 2y 0 2z 0 2 of the worm wheel blank. The centres of the coordinate systems are described with the points O1 and O2 positioned away from each other by the value a. The worm cutting hob rotates around the z 0 1 axis by angle u 0 1, opposite to the trigonometric direction",
            " r \u00f010 \u00de c \u00bc r 1 0\u00f0 \u00de 1 u1; u\u00f0 \u00de \u00f0rectilinear profile\u00de r \u00f010 \u00de c \u00bc r 1 0\u00f0 \u00de 1 u1; h\u00f0 \u00de \u00f0arc profile\u00de \u00f015\u00de The contact lines for the tool of a particular profile are presented in Fig. 7. The generation of the worm wheel tooth flank surface is demonstrated in Fig. 8. In the considerations concerning the determination of the worm wheel mathematical model, the worm model is turned in relation to the z 0 1 axis so that the extreme cutting edge of the tool is located in the central plane, that is, plane y1z1 \u00bc y2z2 (Fig. 1). This position of the worm model is called the basic position. This operation simplifies the presentation of the mathematical model of the worm wheel tooth flank. The parameter u 0 1b is also introduced. It is the angle by which the extreme cutting edge of the machining worm will coincide with the central plane. In the developed algorithm, contact lines are determined. The tool is in the basic position. When solving equation (8), the rotation parameter u 0 1 should include the rotation angle u 0 1b, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002770_j.jmapro.2021.02.015-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002770_j.jmapro.2021.02.015-Figure4-1.png",
        "caption": "Fig. 4. The model of the SLMed circle structure with \u201cskin-core\u201d rotate scanning strategy: (a) layered by \u201cskin-core\u201d; (b) overhead view.",
        "texts": [
            " Differ to the 90\u25e6 rotate scanning strategy, the shape of the external and internal contours fabricated by 67\u25e6 rotate scanning strategy are circle approximately. This is caused by that the 67 is a prime number, the first and last track can occur at any positions when the track number is large enough. In view of the overall situation, the step effect cannot occur and the dimension distribution is nearly uniform. But the shape distortion is still existed in the adjacent layers, as shown in Fig. 3. The model of the SLMed circle structure with \u201cskin-core\u201d scanning strategy is shown in Fig. 4. As can be seen in Fig. 4, the track by the \u201cskin\u201d scanning will cover the external contour and internal contour in every layer. Thus, no matter the rotate angle of the \u201ccore\u201d scanning is, the shape of the external and internal contours are standard circle. The step effect and shape distortion will not existed and the dimension distribution is uniform. Consequently, the SLMed circle structure by the \u201cskin-core\u201d scanning strategy has a better geometric accuracy rather than by the 90\u25e6 and 67\u25e6 rotate scanning strategies. The model of the single-layer was overlaid under the different scanning strategies, the maximum deviation from the designed dimension of each point was extracted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001315_tiv.2019.2955904-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001315_tiv.2019.2955904-Figure4-1.png",
        "caption": "Fig. 4. The clearance between two vehicles in 2.5D environment.",
        "texts": [
            " One might consider approximating the degree 32 curve by a lower degree curve, however, if a single control point of the approximated curve is edited to avoid collision or to decrease curvature to meet the minimum turning radius condition, the curve will no longer lie on the surface. Even the collision-free 2D path is mapped onto the graph surface z = h(u, v), collisions may occur in the path on the 2.5D terrain surface. To avoid such cases, we add two times the margin M to the width of the vehicle when we generate the 2D path. Fig. 4 illustrates the worst case scenario. If we denote the height of the vehicle as \u039b, and referring to Fig. 4, the clearance C between two vehicles can be obtained as: C = 2M cos\u03b2 \u2212 2\u039b sin\u03b2 , (8) which becomes zero when M = \u039b tan\u03b2 . (9) Based on the path c(\u03b6), the motor/engine-torque limitation of the vehicle, and the equation of motion, the time-optimal vehicle control problem can be solved using the algorithm by Bobrow et al. [3] which is studied in Sections V and VI. We define a 3D Cartesian reference coordinate system oxyz, as depicted in Fig. 5. The unit vector along the positive z-axis is given as k"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000235_icpea1.2019.8911194-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000235_icpea1.2019.8911194-Figure3-1.png",
        "caption": "Fig. 3. Mechanical model of THE WTE.",
        "texts": [
            " The fundamental mechanical model can expressed as [3], [10]: dt \u2126d mJgG t =\u0393\u2212\u0393 (2) Where G is the gear speed ratio, \u0393t is the aerodynamic torque, \u0393g electromagnetic torque of the generator or the load torque, ft and fg are the friction coefficients of the turbine and the generator, respectively. The stiffness coefficients of the turbine and the generator (dt and dg) are usually weak can they be neglected. dG and fG are the stiffness and damping coefficient of the gearbox bridge. The equivalent inertia J applied on the high speed shaft is expressed as [14]: J g G J tJ += 2 (3) Figure 3 presents the dynamic model of the implanted wind turbine emulator, the fundamental mechanical equation is giving as [15]: ( ) dt \u2126d mJ gJ DCMgDCM +=\u0393\u2212\u0393 (4) Where \u0393 DCM and \u0393g is the DC motor torque and the generator torque (load torque for DC Motor) respectively. The static model is considered for the representation of aerodynamic power characteristics. They are linked to the dimensions of the turbine and the power coefficient. Fig. 4.a presents the variation of the power coefficient Cp as a function of the speed ratio \u03bb for several given pitch angles \u03b2 [15], [16]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001312_tie.2019.2952780-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001312_tie.2019.2952780-Figure4-1.png",
        "caption": "Fig. 4. Clutch and gearing mechanism of the electromechanical actuator: (a) XY plane 2D representation of the electromechanical actuator, (b) Clutch with its ON-OFF mechanism and (c) rotation of the epicyclic gear with the movement of the clutch.",
        "texts": [
            " The headstock containing the proposed electromechanical actuator is shown in Fig. 3(a). Fig. 3(b) and 3(c) show the isomeric and the YZ plane view of the electromechanical actuator. The electromechanical actuator consists of a clutch, an epicyclic gear and a screw mechanism which converts the rotational motion into a linear motion and performs the clamping and the constant speed spindle rotation integrally with the help of the single drive motor attached to it. The detailed operation of the electromechanical actuator is further explained in the following subsections. Fig. 4 highlights the clutch and the spindle mechanism of the proposed electromechanical actuator. As illustrated in Fig. 4(b), the multi-tooth clutch mechanism is used for the proposed system. Moreover, the motor side plate and the stationary plate attached to the spindle and the gear mechanism also have alternate slot-tooth structure to accommodate the clutch tooth-slot based on clutch operation. The clutch helps to divide the flow of motor power to the clamping or the spindle separately, based on the required operation. Fig. 4(c) shows the epicyclic gear [14] used in the proposed electromechanical actuator model. Two modes of the epicyclic gear are considered for the two different intended operations of the electromechanical actuator, which are explained as follows: 1) Clamping mode: During the clamping mode of the electromechanical actuator, the clutch is coupled with the stationary section as shown in Fig. 4(b) and the ring gear of the epicyclic gear is kept fixed using a spline structure as shown in left side stick diagram of Fig. 4(c), which is called as the planetary arrangement, highlighted in Table I. Furthermore, the spindle is connected with the stationary section and the drawbar, using the spline structure. The clamping mechanism between the spindle and the drawbar is highlighted in Fig. 5(a). Because of this coupling, the arrangement comprising of the clutch, the ring gear, the stationary section, the spindle and the drawbar are restricted from any rotational motion. When the drive motor rotates and the rotational motion is transmitted to the motor side shaft, shown in Fig. 5(a), which makes the sun gear of the epicyclic gear set shown in Fig. 4(c) to rotate at the same speed. This rotational motion is transmitted to the carrier of the epicyclic gear, which rotates at a speed less than that of the sun gear. This is because of the speed reduction using the gear ratio as shown in Table I during the planetary arrangement of the epicyclic gear. This rotation of the carrier is further transmitted to the drawbar, which cannot rotate because of the locking created by the spline mechanism with the spindle, however converts the rotation into linear motion by means of the screw-thread structure of the drawbar and finally generates the clamping force",
            " 6(b), if an tangential force P is acted on the threaded cylinder with a normal force Q, the relation between P and Q can be given as [15]: Q = 1 \u2212 \u00b5stan(\u03b1) tan(\u03b1) + \u00b5s \u00b7 P ; tan(\u03b1) = p 2\u03c0rTC (6) where, \u03b1 is the lead angle, \u00b5s is the coefficient of friction on the screw surface and rTC is the radius of the threaded cylinder, respectively. From (6), the thrust generated by the drawbar can be obtained by introducing Tc and rearranging as follows: P = T rTC (7) F = Tc rTC \u00b7 2\u03c0rTC \u2212 \u00b5sp 2\u03c0rTC\u00b5s + p (8) 2) Spindle mode: During the spindle mode, the clutch is coupled with the motor side plate as illustrated on the right side of Fig. 4(b). This releases the stationary coupling created previously during clamping and the spindle is free to rotate in this mode. However, the thrust produced at the drawbar can still be maintained to its original value. This is because of the self-locking created by the screw-thread mechanism of the drawbar, produced by the friction acting between the screw and threaded cylinder which sustains the produced thrust. During the spindle mode, from the right side stick diagram of Fig. 4(c) where all the parts can rotate, the ring and the sun gear are coupled to rotate together. Thus, when the motor rotates, it will rotate both the sun and the ring gear at the same speed. Moreover, the carrier in between the sun and the ring gear will also rotate with the same speed. This whole process will make the drawbar and the chuck, attached to the carrier to rotate at the same speed as that of the drive motor. This operation, thus called spindle mode is illustrated in Fig. 5(b). Similar to the clamping mode, the locked and the power flow segments for the spindle mode are highlighted with different dashed lines in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002268_012047-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002268_012047-Figure1-1.png",
        "caption": "Figure 1. The airless tire model: Tweel 12N16.5 SSL ALL TERRAIN.",
        "texts": [
            " Series: Materials Science and Engineering 773 (2020) 012047 IOP Publishing doi:10.1088/1757-899X/773/1/012047 integrate into shear band component. It can reduce the complex model and analysis time. The NPT which is developed the shear band can use under the rapid time. Particularly, this NPT model will be advanced in the future. The characteristic of NPT or airless tire, Tweel 12N16.5 SSL ALL TERRAIN, which is developed by Michelin is presented in Table 1. The Tweel airless tire has three main components. There are rubber tread, shear band and polymer spoke (Fig. 1). The tire stiffness testing was performed to investigate the tire stiffness of Tweel airless tire. The airless tire was mounted at the mounting arm of tire stiffness tester (EKTRON TEK model: PL-2003). The compressive force and tire vertical deformation was recorded while the measurement table was moved and compressed the tire. The compressive load of 1000 kg was applied in this study. The experimental result was indicated that the vertical stiffness of airless tire which was studied is 949.37 N/mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003733_ac960909t-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003733_ac960909t-Figure1-1.png",
        "caption": "Figure 1. Electrochemical cell and measurement setup: (A) sensor, (B) holder, (C) Pt electrode, (D) ohmic contact, (E) sample electrolyte (agar film), (F) cation-exchange resin, and (G) proton release.",
        "texts": [
            "2-5 However, in this study, the illumination scan method, which determines the measuring time, was mainly modified to improve the pH imaging as a practical analytical method. The measurement was controlled by a personal computer (PC1:NEC, PC9821Xa). The bias voltage was applied by a customdesigned potentiostat. A semiconductor Laser (Sakai Glass, Tokyo, Japan, \u03bb ) 780 nm of 5 mW maximum power) was used as the light source. During the measurement, the power was set at 10 \u00b5W, and the frequency of the illumination was 5 kHz. Figure 1 shows the design of the electrochemical cell. This electrochemical cell was placed on the X, Y stage during the measurement. The sensor (A) is fixed in a plastic holder (B) so that the holder forms the cell wall. The bias voltage is applied using the Pt electrode (C) and the ohmic contact (D) from the potentiostat. A 27 mm \u00d7 27 mm area of the sensor is in contact with the electrolyte sample (E). In the illumination scanning, the sensor (electrochemical cell) is placed on the two-dimensional X, Y stage with the laser beam fixed",
            " Ion-exchanged water was used for preparing the electrolyte sample. Potassium chloride, sodium hydrate, and agar (Nakalai Tesque, Kyoto, Japan) were of analytical grade. Ionexchange resin (Amberlite IR-120B, sulfonated type, Organo, Tokyo, Japan) was purified by hydrochloride7 unless otherwise indicated. The purified resins were stored in ion-exchanged water. The diameter of the resin was 0.4 mm when they were stored in ion-exchanged water. Measurement Procedure. The measurement procedure is also shown in Figure 1. Agar film was used as the electrolyte, and the cation-exchange resin (F) was placed on it. The gel film was prepared from solution consisting of 1.5% agar and 0.1 M potassium chloride. The solution pH was adjusted to 7.4 by adding sodium hydroxide. The agar film was prepared by pouring heated agar solution directly on the sensor followed by cooling, leaving an agar film 0.5 mm thick on the sensor surface. The cation-exchange resin was placed on the agar film. The transient pH change was imaged, which was generated by the protons released from the resin after penetrating the agar film to the sensor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001955_j.mechmachtheory.2020.104090-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001955_j.mechmachtheory.2020.104090-Figure7-1.png",
        "caption": "Fig. 7. Geometry of the FP and the MP of a semi-regular SPM (SRSPM).",
        "texts": [
            " , 6 , (1) 4 It may be noted that the required time may be reduced further by optimising the code as well as choosing suitable optimal options available in the compiler, which are deemed as parts of future work in this direction. 5 It is mentioned in [25] that the algorithm presented therein is applicable to all architectures; however, no actual implementation is found using that or any other algorithm for these two classes of SPMs. where \u03b8 i is the actuated variable; \u03b3b i , (i = 1 , . . . , 6) , denotes the angular spacing of the line joining the origin to the base of the i thleg, measured in the counter-clockwise direction from the positive X -axis, as shown in Fig. 7 a; and r b is the circumradius of the actual FP of the 6- R SS manipulator, which is taken to be in the form of a semi-regular hexagon in this paper. For the sake of solving the FKP in a manner analogous to the SPM, the fixed and the moving frames of reference are shifted to the centres of the spherical joints located at b \u2032 i and p i , respectively (see Fig. 2 ). Consequently, the points b \u2032 i on the equivalent FP are transformed to: b \u2032\u2032 i = R z ( \u03b3b i )[ r b + l c cos \u03b8i , 0 , l c sin \u03b8i ] \u2212 b \u2032 1 , i = 2 , ",
            " The time taken to solve the FKP by this method, on a computer with specifications same as stated above is about 3.2 ms. The MP, in this case, is a semi-regular hexagon, with alternate sides having equal length. The circum-radius of the platform is denoted by r t . The smaller of the two angles subtended at the centre of the MP by any pair of legs is given by 2 \u03b3t . Assuming that the x -axis of the local frame of reference passes through the point of connection of the platform with the first leg (denoted by t 1 in Fig. 7 b), the angular spacings of the vertices on the MP measured counter-clockwise with respect to the positive x axis are given by: { 0 , 2 \u03b3t , 2 \u03c0/ 3 , 2 \u03c0/ 3 + 2 \u03b3t , 4 \u03c0/ 3 , 4 \u03c0/ 3 + 2 \u03b3t } . The architecture of the semi-regular hexagonal platform has been explained in [2] . For the present example, the circumradius and the angular separation 23 are assumed to be r t = 0 . 5803 and \u03b3t = 0 . 6573 , respectively. The architecture parameters and the input leg lengths have been presented in Table 4 ",
            " It may also be noted that there no improvement in the accuracy, either, as the scalar measure of error, e , is of O(10 \u22128 ) , which is same as that of the previous case (see Table 11 ). The architecture parameters describing a 6- R SS manipulator with semi-regular hexagonal MP as well as FP are discussed in the following. The circum-radius of the FP and MP 24 are assumed to be r b = 1 and r t = 0 . 5803 , respectively. The crank and strut lengths are assumed to be l c = 0 . 5 and l st = 1 , respectively. The position of the connection points with the legs on the FP and the MP are governed by the parameters \u03b3b = 0 . 2985 and \u03b3t = 0 . 6573 , respectively (see Fig. 7 for the geometry of the FP and MP). The input angles are decided upon by performing inverse kinematics (see Appendix B ) for an arbitrarily chosen pose, defined by p x = 0 , p y = 0 , p z = 0 . 8 , c 1 = 0 . 1 , c 2 = 0 . 1 , c 3 = 0 . 2 . The procedure for solving the FKP remains the same as described in Section 3 . The coordinates of the vertices of the moving platform as well as that of the intermediate base have been specified in Table 12 along with the inputs to the rotary actuators. The real solutions to the corresponding FKP have been listed in Table 13 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002162_978-3-030-48977-9_1-Figure1.7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002162_978-3-030-48977-9_1-Figure1.7-1.png",
        "caption": "Fig. 1.7 Example of an electrical vehicle propulsion unit which utilizes a liquid cooled AC\u2013DC converter and 55 kW switched reluctance machine [6, 8]",
        "texts": [
            " The primary design constraint on the volumetric power density of the converter is thermal, i.e., the need to limit operating temperatures and guarantee sufficient thermal cycles of the semiconductor devices and corresponding packages. This implies that the volumetric power density is to a large extend governed by the specific losses of the devices in use, method of cooling, and drive operating conditions. In electric and hybrid vehicles high power density values for machine and converter are essential. An example of such as drive, as shown in Fig. 1.7, utilizes a liquid cooled DC to AC converter with a volumetric power density of 6000 kVA/m3 and a 55 kW switched reluctance machine with a power density of approximately 1.2 kW/kg. The controller and modulator (if used), as shown in Fig. 1.1, are part of an embedded system which is interfaced with the switching device drivers, and sensors (voltage and/or current and position/speed measurements). In addition, these specialized computer systems in the form of digital signal processors or micro-controllers are specifically tailored for electrical drive applications"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001937_ilt-01-2020-0022-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001937_ilt-01-2020-0022-Figure1-1.png",
        "caption": "Figure 1 Geometric model of a textured bearing",
        "texts": [
            " In this work, a cylindrical texture was designed, and the slip on the surface of the bearing was considered. Through an analysis of the slip phenomenon from texturing at the solid\u2013 liquid interface, the coupling effect between texturing and slip was studied, and a distribution of the corresponding combination (denoted \u201ctexture\u2013slip\u201d hereinafter) was designed, which significantly improved the lubrication performance of the bearing. 2. Modelling the surface oil film thickness of a textured bearing Figure 1 shows a geometric model of a textured bearing. In oil film modelling, DH is the thickness variation because of the textured surface,Rp is the dimensionless radius of the cross section,Hp is the Optimal design of the slip\u2013texture Jian Jin, Xiaochao Chen, Yiyang Fu and Yinhui Chang Industrial Lubrication and Tribology maximum depth and c is the dimensionless factor. The depth of a cylindrical texture canbe expressed as follows: DH \u00bc Hp ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi X0 Cx\u00f0 \u00de2 1 Y0 Cy\u00f0 \u00de2 q Rp 0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi X0 Cx\u00f0 \u00de2 1 Y0 Cy\u00f0 \u00de2 q > Rp 8>< >: (1) Considering the influence of slippage, the Reynolds equation is modified with the velocity of the contact surface between the bearing and oil film as the boundary conditions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003542_pime_proc_1948_158_045_02-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003542_pime_proc_1948_158_045_02-Figure15-1.png",
        "caption": "Fig. 15. Diagram of the Axles of a Bogie",
        "texts": [
            " Let the yawing play of each axle relative to the bogie frame be & p ; it is due to the fore-and-aft play of the axleboxes in their guides. Let the lateral play of each axle relative to the frame be & d ; it is made up of the lateral play of the axleboxes Let the wheelbase be 2c, and let d/c = y. in their guides and the end play of the journals in their brasses. - 477c Which &eS, as another condition for no interference by the B Plays thoughout Assume that in their mid-positions the axles are parallel and in track. In Fig. 15 the centre of the front axle is at (XI, y!) and the front wheels make an angle $1 with OX, the centre-he of the track; the cofiesponding quantities for the back axle are x2, yu and $2. The track is assumed to be straight and perfect, and yl, y2, I)~, and t,!t2 are a l l small. Then the lateral displacements of the axles are subject to the conditions KY 1-241) -Y21<2CB +2d Conditions (5) and (6) are the necessary and sufficient conthat is lY1 -y2-2c$ll <2c(p+y) . . . . (la) ditions for no interference by the B and y plays with the sinoidal motions of individual axles moving on straight a d p e r f a and I Y I - Y ~ - ~ C ~ ~ < ~ ~ ( P + ~ ) * "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001967_s11182-020-02124-1-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001967_s11182-020-02124-1-Figure2-1.png",
        "caption": "Fig. 2. Photograph of electric discharge machining: 1 \u2013 metallographic specimen fabrication, 2 \u2013 dog bone specimen fabrication.",
        "texts": [
            " It should be noted that in this table, the current intensity of the electron beam is average, while in the 3D printing process, it is successively changed from 35.0 m\u0410 for the first layers to 28.0 m\u0410 for the upper layers. In order to provide the formation of the melted pool comparable with the required wall width, we use a circular e-beam profile. Using electric discharge machining, dog bone and metallographic specimens were cut from the EBAM vertical walls to conduct the tensile strength tests and metallographic measurements, respectively. This is illustrated in Fig. 2. The surface structure was investigated using an Altami MET 1C metallographic microscope (Russia) after its polishing by abrasive machining with the different sandpaper grit, diamond polishing paste with a soft tissue and then chemical etching for 40 seconds in Kroll\u2019s reagent. The microhardness testing was performed on a Duramin 5 (Denmark) hardness tester at a 100 g indentation load. Static tension tests of the dog bone specimens were carried out on a universal testing machine UTS-110M-100 (Russia)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003907_nme.1620380611-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003907_nme.1620380611-Figure1-1.png",
        "caption": "Figure 1. The rod geometry",
        "texts": [
            " For these reasons, formulations comprising more general cases are necessary to treat the problem from a mechanically less restricted point of view. A most efficient solution can actually be achieved by defining the element stiffness matrix which originates from the true governing equations of the bar. For this purpose, the appropriate set of the governing differential equations must be solved. 2. FIELD EQUATIONS FOR STATICAL ANALYSIS O F THE SPATIAL BAR 2.1. The rod geometry Consider a naturally curved and twisted rod. The trajectory of the geometric center G of the rod is defined as the rod axis (Figure 1). The vectors t, n, and b are the tangential, normal and SPATIAL BARS 1033 binormal unit vectors, respectively. The relationships among these unit vectors are defined by the Frenet formulas, which are rewritten as dt/ds = xn, dn/ds = zb - Xt, db/ds = - t n (1) It is not that x is always positive, and t is positive for a clockwise rotation about t when advanced in the increasing s-direction. For planar rods T = 0, and for straight rods t = x = 0. In order to take into account the initial (pre-) twist of the cross-section, a second rectangular Cartesian frame is defined such that the x,-axis is the direction oft , and the xz-, x3-axis are the principal axes of the cross-section. From Figure 1, t = i l n = iz cos 8 - i, sin 8 b = it sin 8 + i, cos 8 (2) With the aid of equations (1) and (2), the differential relations among the unit vectors (il,i2,i3) are dip _ - ds - & p j k X k i j where (4) x1 = t +-, x 2 = Xsin8, x3 = ~ c o s 8 and E~~~ is the permutation tensor. It is observed that even for the simplest case of a straight rod, x1 will not be zero in the presence of an initial twist. This is the case for perforator, propellar and turbine blades. d8 ds 2.2. The equations of geometric"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001993_s00170-020-06278-7-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001993_s00170-020-06278-7-Figure1-1.png",
        "caption": "Fig. 1 Two styles of flute profile. a \u201cOpen\u201d flute. b \u201cClosed\u201d flute",
        "texts": [
            " The remainder of this paper is organized as follows: the problems in two-pass grinding of tool flute are given in Section 2, followed by detailed methods and their models to control grinding wheel positions and orientations in Section 3. To validate the proposed approach, a set of numerical simulations and experiments are presented in Section 4. Finally, Section 5 concludes this paper, together with the planned future work. There are two styles of the flute profiles, \u201copen\u201d and \u201cclosed,\u201d and the major difference in terms of profile parameters. As shown in Fig. 1a, there is no inflection point in the tail curve of \u201copen\u201d flute. Therefore, the \u201copen\u201d flute can be represented by three parameters, inner core radius Rc, rake angle \u03b3o, and edge width Ew.As shown in Fig. 1b, there is an inflection point in the tail curve of \u201cclosed\u201d flute. Except for inner core radius Rc, rake angle \u03b3o, and edge width Ew, the outer core radius Rcb is needed for describing the \u201cclosed\u201d flute. The \u201cclosed\u201d flute profile is usually used in two edge cutting tool. The difference of grinding processes between an \u201copen\u201d flute and a \u201cclosed\u201d flute is reflected in the contact zone between wheel and tool. In \u201copen\u201d flute grinding, the wheel is involved in grinding partially in width direction; however, the whole grinding wheel cylindrical surface totally participates in material removal of \u201cclosed\u201d flute cutting, ,as shown in (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003542_pime_proc_1948_158_045_02-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003542_pime_proc_1948_158_045_02-Figure2-1.png",
        "caption": "Fig. 2. Cross-sections of Full-size Tyres and Rail Heads",
        "texts": [
            " The motion of a real axle is of course much affected by the imperfections of the track, which constantly bring the flanges, and probably also the axlebox guides, into play. Mauzin (1933) found that high or low spots in one rail had a pronounced effect on the lateral oscillation. None the less, full-size records suggest that the natural motion of a single axle may well be the fundamental factor; it is hard to overlook the general agreement in wavelength. At any rate, there seems to be ample justification for attacking the problem in its simplest formthe motion of a single axle, particularly when the primary object is to study flange-rail impacts. Fig. 2a shows the profile of a new tyre on a new rail, the rail being shown in the two extreme positions relative to the wheel which are permitted by the flange clearance of 3 inch. The corresponding points of treadrail contact are A and B, and the effective tread radii are ra and 4. Clearly, when one wheel of a pair is running on radius r. Tyre and Rail Profiles. at WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from THE MOTION OF the other is running on radius rb, and the maximum difference between the ef\u20acective tread radii is (rp-rb) = 3\\20 = 0.038 inch, assuming a coning angle of 1 in 20. Fig. 2b shows a worn tyre on a worn rail; the profiles were selected from a number supplied by the London, Midland and Scottish Railway Company, and seem to be representative. Whereas with new profiles the position of the point of contact on the rail hardly moved, with these worn profiles, owing to the concavity of the tread, it moves nearly 2 inches between its extreme positions C and D. The point of contact on the tread therefore moves more than 23 inches laterally, and the maximum difference between the effective tread radii (rc-rd), is much greater than (ra-rb)",
            " It follows that, for hunting to persist, the axle must yaw during the period of flange contact so that the angle of incidence is destroyed and a definite angle of reflection is built up; this is true whether flange impact is due to Mturd oscillation or to track imperfections. The angle of yaw during flange contact depends on the rate of yaw in radians per unit distance travelled and on the distance for which flange contact lasts. The rate of yaw is, by simple geometry, h/2br, where h is the difference between the effective radii of the wheels. The effective radius of the south wheel is its least mead radius (ra or rd of Fig. 2) ; that of the north wheel is greater than its greatest tread radius, since t h i s wheel is running partly on its flange in the case of two-point contact, or on the root of its flange if there is single-point contact (Fig. 2). Thus h may well exceed 4 inch. In the absence of inertia forces, lateral friction at the treads maintains flange contact while $ is positive, and causes separation when # becomes*negative. Flange contact therefore ceases when the axle is square with the track; pure roiling is then resumed at the treads and the north wheel leaves the rail tangentially. So at low speed the flanges steady the motion, because, whatever the angle of incidence, the angle of reflection is zero. At high speed the lateral momentum may be d c i e n t to maintain flange contact after # has become negative, and thus allow an angle of reflection to be built up",
            " Since wear can only occur where there is contact, it must tend to ptoduce, and then maintain, tyre and rail profiles such that the whole width of the rail surface is used. Further, since the tyre is wider than the rail, the tyre profile becomes concave and the rail profile remains convex. Owing to the lateral motion of the axle, such profiles cannot fit exactly in any one position, but they will tend to be such that, as the axle traverses across the track, the contact areas traverse across the width of the rail surfaces in the opposite direction, as shown in Fig. 2b. If the proportions of new, part-worn, and badly worn tyres in operation do not vary much from time to time, it may be expected that as a rail wears down it soon reaches a definite profile (generated by the \u201caverage tyre\u201d), which profile it thereafter retains. This profile will not be constant from place to place, because the conditions of wear will be affected by curves, small errors of gauge, etc. (similarly, one would expect the minor imperfections of the vehicles to cause their wheels to wear to slightly different profiles)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002535_s00170-020-06330-6-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002535_s00170-020-06330-6-Figure5-1.png",
        "caption": "Fig. 5 Spatial position relation of wide-beam laser-powder flow",
        "texts": [
            " The powder fluxC (g/(mm2*s)) can be obtained by[18, 20] C \u00bc Cmaxe \u2212 2y2p r2 zp\u00f0 \u00de \u00f02\u00de where Cmax is the maximum powder flux at the center, r(zp) is the effective radius at a distance zp from the nozzle outlet, and Cmax satisfies the following equation: qm wp \u00bc \u222b r zp\u00f0 \u00de \u2212r zp\u00f0 \u00deCmaxe \u2212 2y2p r2 zp\u00f0 \u00dedyp \u00f03\u00de where qm (g/s) is the powder feed rate and wp is the width of the powder beam in the x direction. According to the powder flux, the powder particle number concentration n(1/mm3) can be obtained by n xp; yp; zp \u00bc Cmax vpmp e \u2212 2y2p r2 zp\u00f0 \u00de \u00f04\u00de where vp (mm/s) is the particle velocity along its central line, namely, zp axis, and mp (g) is the average mass of the particle mp \u00bc 4 3 \u03c0r3p\u03c1p \u00f05\u00de where rp (mm) is the particle radius and \u03c1p is the particle density, g/mm3. In Eq. (4), r(zp) is the effective radius of the powder beam r zp \u00bc r0\u2212zptan \u03b8 \u00f06\u00de Figure 5 shows the geometric position relationship between the laser beam and the powder beam. And the powder beam axis has an inclined angle \u03c6 and height H with respect to the substrate. According to the relationship between the powder beam coordinate system Opxpypzp and the laser beam coordinate system Oxyz, as shown in Fig. 5, the coordinate transformation can be obtained: xp \u00bc x yp \u00bc y\u2212H=tan \u03c6\u00f0 \u00desin \u03c6\u2212 z\u2212H\u00f0 \u00decos \u03c6 zp \u00bc y\u2212H=tan \u03c6\u00f0 \u00decos \u03c6\u00fe z\u2212H\u00f0 \u00desin \u03c6 \u00f07\u00de Then the powder particle concentration distribution n(xp, yp, zp) in the laser coordinate system is n x; y; z\u00f0 \u00de \u00bc Cmax vpmp e \u22122 y\u2212H=tan \u03c6\u00f0 \u00desin \u03c6\u2212 z\u2212H\u00f0 \u00decos \u03c6\u00bd 2 r2 z\u00f0 \u00de \u00f08\u00de where r (z) can be expressed as r z\u00f0 \u00de \u00bc r0\u2212 y\u2212H=tan \u03c6\u00f0 \u00decos \u03c6\u2212 z\u2212H\u00f0 \u00desin \u03c6\u00bd tan \u03b8 \u00f09\u00de When the laser (Iz) acts on\u0394z distance powder flow, according to the principle of Mie and Lambert-Beer\u2019s law [9], the power density attenuation dI can be calculated by dI \u00bc \u2212kext\u03c0r2pI x; y; z\u00f0 \u00den x; y; z\u00f0 \u00dedz \u00f010\u00de where kext is the extinction coefficient, and the attenuated laser power density IA can be obtained by IA x; y; z\u00f0 \u00de \u00bc I x; y; z\u00f0 \u00deexp \u2212kext\u03c0r2p \u222b zp x;y\u00f0 \u00de z n x; y; z\u00f0 \u00dedz \" # \u00f011\u00de The attenuation coefficient \u03b7A is defined as \u03b7A \u00bc exp \u2212kext\u03c0r2p \u222b zp x;y\u00f0 \u00de z n x; y; z\u00f0 \u00dedz \" # \u00f012\u00de where zp(x, y) is the upper interface between the powder beam and the laser beam. To simplify the calculation, zp(x, y) is defined as the straight line PQ (as shown in Fig. 5), which is calculated by [9]. zp y\u00f0 \u00de \u00bc K \u00fe y\u2212K=tan\u03c6\u00f0 \u00de tan \u03c6\u2212\u03b8\u00f0 \u00de \u00f013\u00de where K = H + r0sin\u03c6/tan\u03b8, r0 = 0.5 mm. When the powder particles pass through the laser beam, its temperature will increase. The temperature increment \u0394T at time interval \u0394t can be calculated by the following equation [9, 18]: \u03b1IA x; y; z\u00f0 \u00de\u03c0r2p\u0394t \u00bc 4 3 \u03c0r3pcp\u03c1p\u0394T \u00f014\u00de where \u03b1 is the laser absorption for the powder particle, \u0394t = dz/vpsin\u03c6 is the particle flying time. The powder particle temperature can be calculated by the following equation [9, 18]: Tp x; y; z\u00f0 \u00de \u00bc T0 \u00fe 3\u03b1 4cp\u03c1prpvpsin \u03c6 \u222b zp \u2212y\u00f0 \u00de z IA x; y; z\u00f0 \u00dedz \u00f015\u00de where T0 is the ambient temperature",
            " In the following calculation, the Fe-based alloy powder is used as the powder material. Table 1 lists the main parameters used in the calculation of the laser power interaction model. Figure 6a shows the attenuation coefficient distribution of the wide-beam laser attenuated by the powder cloud. As can be seen from the Fig. 6a, the attenuation coefficient decreases along the +y axis, which indicates that more laser energy attenuated by the powder in the +y direction. This can be explained as follow; as shown in Fig. 5, the laser-powder interaction zone in +y direction is larger than that in \u2013y direction, which means that more laser power is attenuated on the right side of the y axis. Figure 6a also shows that the minimum value of attenuation coefficient is 96.56%, which means that the maximum attenuation of laser power is 3.44%. Loss of the laser power is absorbed and scattered by the powder. And this calculation results are close to those in reference [18]. In their study, the laser power attenuation ranged from 0",
            " This is because the powder beam is located on the right side of the laser beam (+y direction). And the powder temperature calculation results are also close to the reference [18], and they calculated that the maximum powder temperature ranges from 62.3 to 172.1 \u00b0C. In addition, the particle temperature in the positive y-axis is slightly lower than that in the negative y-axis. This is due to the inconsistent flying time of the powder particles in the different zone of the laser beam. As shown in Fig. 5, the particle moves from point P (+y direction) to point Q (\u2212y direction), and it has the longer flying time, so the powder temperature on the left side of y-axis is higher. From the previous section, it can be found that there are many parameters affecting the laser-powder interaction. In order to reveal the influence of these parameters on the laser energy attenuation coefficient and the powder particle temperature distribution, 7 parameters, namely, as laser power (P), powder feeding rate (qm), particle velocity (vp), mean particle radius (rp), laser-powder angle (\u03c6), powder divergence angle (\u03b8), and nozzle outlet height (H), are selected"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000158_j.mechmachtheory.2019.103625-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000158_j.mechmachtheory.2019.103625-Figure10-1.png",
        "caption": "Fig. 10. CAD model of M3, showing together the coupler curves generated by the mechanism, (a) one branch of the coupler curve, where all prescribed positions are located, (b) the other branch.",
        "texts": [
            "495948] [0.435518, \u22120.700771] angles for this linkage in visiting 10 poses, which shows clearly a monotonically changing input angle. Fig. 9 b shows the coupler curve of the linkage synthesized. It is seen that the curve has two branches, with one of them all poses are located. It is noted that the curve plot shows disconnected points, which is due to a resolution issue of the intersection function in Maple. We further verify the feasibility of the mechanism in SolidWorks motion study, as shown in Fig. 10 , where the coupler curves traced by the mechanism are clearly displayed. The motion study confirms that all predefined positions are contained in one branch, as displayed in Fig. 10 a. Moreover, the motion study shows all poses are visited in the right order. The other branch of the coupler curve in 2D is shown in Fig. 10 b for information. In IO analysis, we first take a look at the linkage of Example I. Using the data from Table 2 , lengths of each link in the mechanism can be calculated. This allows further to find the output angle/position for any input. The equation yields two real solutions, as given in Table 5 . While the IO equation admits maximum up to six real solutions, this linkage has only two real solutions for the given input. We include one more example of IO analysis for a linkage, noted as M4. The linkage has three limbs of equal length with 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003672_ieeestd.2000.91147-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003672_ieeestd.2000.91147-Figure7-1.png",
        "caption": "Figure 7\u2014R-X diagram",
        "texts": [
            " In addition, protection characteristics can be shown on time-current diagrams, R-X diagrams, and one-line diagrams with zones or time characteristics overlaying the lines. Figure 6 shows one method of representing protective relay characteristics. This particular characteristic is the current versus time characteristic of a time overcurrent relay. The characteristic shows two regions: one in which the relay operates, and the other in which the relay does not operate. The line separating the regions is the characteristic curve of the relay. Figure 7 shows another method of representing relay operating characteristics using an R-X diagram. Again, the characteristic separates two regions. Authorized licensed use limited to: ULAKBIM UASL - Anadolu Universitesi. Downloaded on May 01,2014 at 23:43:40 UTC from IEEE Xplore. Restrictions apply. IEEE Std C37.113-1999 IEEE GUIDE FOR PROTECTIVE RELAY 10 Copyright \u00a9 2000 IEEE. All rights reserved. The selection of line protection requires the consideration of several factors, some of which are mutually exclusive"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001166_j.procir.2019.03.197-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001166_j.procir.2019.03.197-Figure2-1.png",
        "caption": "Fig. 2. Differences in DOF between as-built and installed assemblies.",
        "texts": [
            " As it is well known, the degree of freedom (DOF) of a whole assembly results from the sum of the remaining DOFs per part, which are not locked by the adjusted joints and the bearings. In order to guarantee the mobility of mechanisms they are usually kinematically underconstrained. Moreover, additively manufactured non-assembly mechanisms can have more DOFs than in installed state depending on the number of bearings to be included in the manufactured assembly. This fact is examplarily shown for a five bar linkage mechanism in Fig. 2. Assuming that the frame including three bearings with one rotational DOF, the DOF of the as-built assembly increases from one to four compared to the installed state of the mechanism. The decision variables vector x for the optimzation includes all possible translations and rotations of the individual parts vi,red, which are not locked by a joint or a bearing: x = [ vi,red, . . . , vn,red]T. (2) Depending on the objective in focus, the optimization problem can be formulated as a constrained single-objective optimization problem in order to minimize the build time T or rather the support material quantity Q: min ob j = T ( x) or ob j = Q( x), (3) subject to Vi \u2229 Vj = \u2205 \u2200 i, j = 1 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001265_s12555-019-0101-x-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001265_s12555-019-0101-x-Figure5-1.png",
        "caption": "Fig. 5. Line of sight tracking.",
        "texts": [
            " Mapping to attitude commands Because of the underactuated characteristic, the motions of the quadrotor along the x and y axes are related to its attitude angles. So in order to control the motion along x or y, we should control the pitch angle \u03b8 and roll angle \u03d5 . Inverse mapping method is used here to obtain the commands of roll angle \u03d5 , pitch angle \u03b8 and lift U1 from the virtual acceleration signals. We assume that the quadrotor is oriented towards the target before and during the standoff process, which leads to a line of sight tracking. According to the geometry shown in Fig. 5, the desired yaw angle can be specified as \u03c8d = \u03c6 +\u03c0. (19) From the relationship between the x\u2212y and \u03c1 \u2212\u03c6 coordinates, we have a\u03c1 = ax cos\u03c6 +ay sin\u03c6, a\u03c6 =\u2212ax sin\u03c6 +ay cos\u03c6, (20) where ax and ay are accelerations along x and y axes respectively. Note that the first three equations in (3) can be rewritten as max =U1 (cos\u03c8d sin\u03b8d cos\u03d5d + sin\u03c8d sin\u03d5d) , may =U1 (sin\u03c8d sin\u03b8d cos\u03d5d \u2212 cos\u03c8d sin\u03d5d) , maz =U1 cos\u03b8d cos\u03d5d \u2212mg. (21) Substituting (19) and (20) into (21), we get sin\u03b8d cos\u03d5d =\u2212 m U1 a\u03c1 , sin\u03d5d = m U1 a\u03c6 , cos\u03b8d cos\u03d5d = m U1 (az +g)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001160_codit.2019.8820315-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001160_codit.2019.8820315-Figure1-1.png",
        "caption": "Figure 1. Delta wing aircraft axis",
        "texts": [],
        "surrounding_texts": [
            "two nonlinear control schemes to control the angle of roll channel for delta wing aircraft with the presence of wing rock phenomenon. The proposed nonlinear controllers for the considered aircraft are the sliding mode backstepping controller and disturbance observer-based sliding mode backstepping controller. The stability of the roll-channel system based on both nonlinear controllers is analyzed using Lyapunov method and the zero-convergence of roll angle error and estimation error of disturbance observer has been guaranteed and proved. The performance of the proposed controllers has assessed and compared in terms of transient characteristics and the robustness characteristics. The effectiveness of the controllers for the roll-channel system of aircraft has been verified via computer simulation within the environment of MATLAB package.\nKeywords: Delta wing aircrafts, wing rock, backstepping, sliding mode backstepping, disturbance observer.\nI. INTRODUCTION\nDelta wing aircrafts development is highly funded aircrafts projects. Nowadays, many types of such aircrafts are still in military and commercial usage. The delta wing aircraft is characterized by high level of speed ranges and maneuverability. Its dynamic is complex and highly nonlinear [1]. The rolling motion of the delta wing aircraft are the rotation of the aircraft on the rolling axis. Rolling axis is the line passing through the aircraft\u2019s fuselage and the center of gravity as indicated in Figure (1). The ailerons at the end of the wing are the movable parts responsible for rolling motion control of the aircraft [2].\nWing rock phenomenon is an undesired motion appears in high angles of attack. This phenomenon comes out due to the presence of a limit cycle caused by wing rock dynamics. The wing rock phenomenon results in rolling the aircraft in both direction of roll angles with specified amplitude and frequency. The dynamics of wing rock phenomenon can be seen in many researches and studies [3-5].\nIn what follows, some recent relevant control methodologies for wing rock phenomenon are briefly interviewed; Finite form adaptation has been proposed by (Lee and et.al, 2016) [6]. Also, Adaptive PID controller using selfrecurrent wavelet neural network identifier are presented by (Malekzadeh, et.al, 2016) [7]. Robust MRAC using novel weighted sigma modification has been proposed by (Humaidi and etal, 2017) as indicated in [8]. Robust Backstepping Control Using Disturbance Observer has been performed by (Wu, et.al , 2017) in [9]. (Cho, etal, 2018) developed Analytic Solution of Continuous-Time Algebraic Riccati Equation for Two-Dimensional Systems [10]. A singular perturbation approach to saturated controller design has been proposed by (Rayguru and Kar, 2018) [11]. In the above literature survey, most of the proposed techniques are either individual controllers or hybrid controllers, which are mixing two or more control strategies to gain the benefits of the composite control strategies such as to compensate the weakness of the others like the case of the present work.\nBackstepping control algorithm is a powerful procedural control strategy. It is established according to iterative steps which ends when the control action reaches the channel of desired state. Throughout the design procedure of backstepping control algorithm, virtual controllers are assigned to intermediate state variables [12]. However, the design of this procedural algorithm requires a prior knowledge of both systems uncertainty and parameters. As such, the concept of convergence in the presence of uncertainty, like an exertion of disturbance, will be different from that in case of no uncertainty. [13].\nSliding mode control (SMC) is a nonlinear discontinuous control strategy, which varies the structure of dynamic system by application of set-valued control signal such as to enforce the system trajectory to slide on a surface, which defines the normal behavior of the system. Therefore, the existence of sliding surface is the milestone in designing SMC. Whenever the trajectory of error state is enforced to reach this sliding\n978-1-7281-0521-5/19/$31.00 \u00a92019 IEEE\n-1215-",
            "surface, it asymptotically slide to equilibrium point. The SMC has good robust characteristic against uncertainty in system parameters and also against external disturbance with bounded matched property [14]. However, the robustness characteristics against uncertainty in parameter is restricted due to the pre-requisition of nominal value and upper bound of this uncertainty [15, 16]. The robustness capabilities may be altered when dealing with totally unknown disturbance bound.\nThe nonlinear disturbance observer is an efficient tool to estimate the unmeasured disturbance. Including the estimated disturbance in the control law will grant the control law the advantage to cope with unperturbed system and compensate the uncertainty in an efficient manner [17, 18].\nFurthermore, combining the features of sliding mode control with backstepping control and the disturbance observer will acquire the benefits of performance characteristics and robustness of the three strategies and yield what so called the Sliding mode backstepping control with disturbance observer.\nIn the present work, the performance of two controllers named sliding mode backstepping controller and sliding mode backstepping controller with disturbance observer are investigated for wing rock phenomena in the roll dynamics of delta wing aircraft.\nThe present work of the paper is motivated and encouraged by the recent studies in robust and nonlinear adaptive control design of different applications presented by (Humaidi and Hameed) [8, 16, 19].\nThe development of two adaptive control strategies to control angular position of the roll-channel system for delta wing airplane is the essential contribution of this paper in order to fulfill the following tasks:\n1. To robustly control the roll angle of a delta wing aircraft in presence of wing rock phenomena using sliding mode backstepping control, 2. To robustly control the roll angle of a delta wing aircraft in presence of wing rock phenomena using sliding mode backstepping control with disturbance observer, 3. To cope the unknown (upper bounded) exerted disturbance using sliding mode backstepping control, where disturbance upper bound is needed, 4. To cope the unknown upper bound of exerted disturbance using sliding mode backstepping control with disturbance observer, where disturbance is estimated, also the disturbance observer can be employed as a chattering suppression tool. 5. Dealing with constrained control effort represented by the limited ailerons deflection which is a limitation avoided by almost all the researchers.\nThe whole paper is divided into the following sections. In section two, the dynamic model of the airplane with wing rock phenomena is developed, Section three presents the control design and stability analysis based on Lyapunov method, section four shows the effectiveness of both control strategies and make a comparison study via simulation results and finally Section five highlights the main concluded points.\nWing rock phenomenon were studied and analyzed by many researchers in the literature. The derived model is\nformulated by fitting the experimental data gathered from wind tunnel simulation to a mathematical expression agreed to describe the wing rock phenomenon. The models got from the experimental fitting are a second order nonlinear models differs from each other\u2019s by nonlinearity terms. One of the first most cited works for modern control theory is the model proposed by [3]. A further development on the model is suggested by [4]. The most recent wing rock phenomenon suppression controllers were designed considering the result of [5] depending on the original work in [20]. For a case of one degree of freedom which means that the considered movement in the aircraft in the air tunnel is the roll motion as in Figure (2).\nThe experimental procedure has done for an 80\u00b0 delta wing aircraft with a number of aircrafts configuration and for different angle of attack simulation cycles for each configuration as in Figure (3).\nIn this paper the C configuration in the Figure (3) will be considered as a case study for the proposed controllers. The C configuration is the closest configuration for the delta wings aircrafts in use either fighter or civilian aircrafts. The C configuration is consist the delta wing (80\u00b0swept wing) and the forebody and the nose tip. The differential equation that describe the C configuration is constructed as:\nCoDIT\u201919 | Paris, France - April 23-26, 2019 -1216-"
        ]
    },
    {
        "image_filename": "designv11_14_0002110_nap51477.2020.9309696-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002110_nap51477.2020.9309696-Figure1-1.png",
        "caption": "Fig. 1. Models (samples) that were built in Visual-Environment: a - hollow cylinder, b - cone, c - hollow triangle, d - hollow prism.",
        "texts": [
            " Goldak [17] with normal (Gaussian) distribution of specific heat along all coordinate axes in the heat flux with a shape of an ellipsoid is moslty used for analysis of thermal processes of arc welding and cladding. The main feature of this model is the independent setting of the distribution of specific heat output. Calibration of the model by J. Goldak remains the most difficult issue of its direct application for a purely theoretical analysis of thermal welding processes. The CEA software called \"Visual-Environment\" was used to study the formation of the residual SST in the printed prototypes of products of various shapes. The following models (samples) were considered (Fig. 1): a hollow cylinder with a diameter of 40-150 mm, a cone with a base diameter of 60 mm, a hollow triangle with a side of 50 mm with radius of the angle section of 5-20 mm and a hollow prism with a side of 50 mm and with radius of curvature of the trajectory at an angle from 5 to 20 mm. Experiments were performed using 10 mm thick E 235-C steel plates, the chemical composition of which is shown in Table I. Layer-by-layer cladding was performed with a G4Si1 electrode wire of solid cross-section with a diameter of 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001978_0954407020964625-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001978_0954407020964625-Figure6-1.png",
        "caption": "Figure 6. Detailed fault gear diagrams.",
        "texts": [
            " Fault diagnosis cases under variable rotational speed Case 1: fault diagnosis of a planetary gearbox Data description. A planetary gearbox fault implantation bench is adopted for signal acquisition under variable rotational speed. As exhibited in Figure 5, the test bench contains a motor, a planetary gearbox, two bearing seats and two shaft couplings. Three planet wheels are designed in the gearbox. The teeth numbers of the sun wheel and planet wheels are 36 and 18, and the module n is 1.5, respectively. There are ten health conditions as shown in Figure 6: normal condition (NC); three sun wheel fault types (crack, pit and worn tooth), which are named as WC, WP and WW; three pinion fault types (crack, pit, and worn tooth), which are named as PC, PP and PW; and three coupled fault types (wheel worn and pinion worn, wheel pit and pinion crack, wheel pit and pinion worn), which are named as WWPW, WPPC and WPPW. The data acquisition system is LMS SCADAS, and the involved software is LMS Test.Lab. The acceleration sensor is DH311E made by Donghua Inc., which is installed on the upper surface of the gearbox at a sampling frequency of 12"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001662_j.addma.2020.101294-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001662_j.addma.2020.101294-Figure8-1.png",
        "caption": "Figure 8 Determination of the voxel size based on Rosenthal's equation a) Simulated melt pool shape by using Rosenthal's equation, b)Voxel size created based on the melt pool dimensions (width, length, depth)",
        "texts": [
            " The voxel width, length and depth were determined based on the melt pool dimensions in steady state, as approximated by using Rosenthal\u2019s equation (Equation (1)), with a reference to ambient temperature. The equation is resolved in 3D to simulate the temperature field, with the coordinate system located at the surface and instantaneous beam location. The melt pool dimensions are approximated based on the boundary where the temperature reaches the melting temperature of the material; the melting boundary contour is therefore assumed to be the boundary of the melt pool. The voxel xyz dimensions are approximated by using the melt pool dimensions (width, length and depth), which can be seen in Figure 8 for a Ti64 material system under laser power 410 W and scan speed 8.5 mm/s. Based on these parameters, the voxel size is set to 1.6 mm in length, 1.3 mm in width and 0.8 mm in depth. The voxel size was kept constant through the simulation for each Jo ur na l P re - ro of 18 combination of process parameters. In future work, it can be made to change dynamically based on the dynamic changes in process parameters and substrate temperature. The limitation in the model proposed by Li et al. [29] is that it can be applicable to keeping track of the thermal history only for straight lines (1D toolpath)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure25.2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure25.2-1.png",
        "caption": "Fig. 25.2 Schemeof the formation of the bendof the link:a acting forces,bgeometry for calculating the value of the shoulder action of the force Fi",
        "texts": [
            " This approach is presented in this paper,where the author proposed a kinematicmodel of the manipulator taking into account the direction of action of the forces of gravity and friction between the elements in the links and also describes the mechanism of bending of the link in its entire range. The manipulator under consideration (see Fig. 25.1a) is a set of series-connected links, each of which consists of solid-state elements with a spherical surface in the form of circular disks of a given thickness (see Fig. 25.1b, c). Elements are characterized by the followingparameters (seeFig. 25.2b):R is the radius of curvature of the surfaces of the elements by which they are in contact with each other; H is the height of the cylindrical part of the element; D is the diameter of the element; d is the diameter of the circle on which the holes for the cables lie; P is the weight of the element. Through the holes on the edges of the elements are spring-loaded flexible cables, the ends of which are fixed to the last elements of the links. Springs provide the necessary tension of the cables and are elements of the displacement and tension sensors of the cables, combined in one unit with the drives (based on stepper motors with gearboxes) (see Fig",
            " The components of the Rot j,i (l j , \u03b2 j i ) matrix are calculated by the following formulas: Rot11j i = sin2 \u03c8 j (1 \u2212 cos\u03b2 j i ) + cos\u03b2 j i , Rot12j i = \u2212 sin\u03c8 j cos\u03c8 j (1 \u2212 cos\u03b2 j i ), Rot13j i = \u2212 cos\u03c8 j sin \u03b2 j i , Rot21j i = \u2212 sin\u03c8 j cos\u03c8 j ( 1 \u2212 cos\u03b2 j i ) , Rot22j i = cos2 \u03c8 j (1 \u2212 cos\u03b2 j i ) + cos\u03b2 j i , Rot23j i = \u2212 sin\u03c8 j sin \u03b2 j i , Rot31j i = cos\u03c8 j sin \u03b2 j i , Rot32j i = sin\u03c8 j sin \u03b2 j i , Rot33j i = cos\u03b2 j i . The components of vector p j i (\u03b1 j i , \u03c8 j ) are defined as: pxji = (\u22122R sin \u03b1 j i + h sin 2\u03b1 j i ) cos\u03c8 j , py ji = (\u22122R sin \u03b1 j i + h sin 2\u03b1 j i ) sin\u03c8 j , pzji = \u22122R cos\u03b1 j i + h(cos 2\u03b1 j i + 1). The constant h is the distance from the center of curvature of the surface of the element to its center of coordinates (see Fig. 25.2b). In accordance with this model, the position vector of the center of the 1st element of the 1st link is calculated as follows: p11 = T11pT 0 , where pT 0 = [0; 0; 0; 1]T is the position vector of the center of the base fixed element. The last element of each link is a reference for the next link, and the position vector of the center of the 1st element of the jth link is p j1 = T j ipT ( j\u22121)M( j\u22121) . The position vector of the center of the ith element of the jth link is p j i = T j ipT j(i\u22121). Accordingly, the position vector of the operating point of the manipulator is the position vector of the center of the last element of the last link: pt = pNMN . Consider the formation of the bend of the 1st link, taking into account all the forces acting on it. Themovable element 1 is affected by: the force of gravityP (weight of the element), normal reaction force N1 arising at the contact pointMc1 of elements 1 and 2, friction force of adhesion Fs1 and force F2 (see Fig. 25.2a). For the case of static equilibrium of the moving element 1 relative to the stationary one, for some rolling angle \u03b11, we write down the equations of equilibrium of moments and projections of forces on the selected axes (action vectors of forces N1 and Fs1): N1 + P cos\u03b11 \u2212 F2 cos(\u03b11 + \u03b12) = 0, (25.3) \u2212P sin \u03b11 \u2212 F2 sin(\u03b11 + \u03b12) + Fs1 = 0, (25.4) fr N1 \u2213 PdP1 \u2212 F2dF2 = 0. (25.5) Here fr is the coefficient of rolling friction. The sign of the moment created by gravity P in Eq. (25.5) depends on its direction relative to the contact point of the elements",
            "5): N1 = F2 cos(\u03b11 + \u03b12) \u2212 P cos\u03b11, F2 = (P( fs cos\u03b11 + sin \u03b11))/( fs cos(\u03b11 + \u03b12) \u2212 sin(\u03b11 + \u03b12)), PB1( fr cos(\u03b11 + \u03b12) \u2212 dF2)/( fs cos(\u03b11 + \u03b12) \u2212 sin(\u03b11 + \u03b12)) + PA1 = 0, where PB1 = P( fs cos\u03b11 + sin \u03b11); PA1 = \u2212P( fr cos\u03b11 \u00b1 dP1). The calculation of the shoulder of the action of gravity dP1 acting on element 1 is made through the absolute coordinates of the contact point of elements Mc1 and the center of the coordinate system O1 of element 1, which, in turn, are determined through transformationmatrices (25.2) for a given angle \u03b11. For the remaining elements, the calculation of dPi is similar. The calculation of the dF2 value is carried out geometrically (see Fig. 25.2b) and for all pairs of contacting link elements the same: dFi = Li\u22121 sin(\u03b1i \u2212 \u03b1\u2032 i\u22121), Li\u22121 = \u221a (R\u2032 i\u22121) 2 + (R sin \u03b1i\u22121)2, \u03b1\u2032 i\u22121 = tan\u22121(R sin \u03b1i\u22121/R \u2032 i\u22121), R \u2032 i\u22121 = 2 \u221a (R2 \u2212 (D/2)2) \u2212 R cos\u03b1i\u22121 \u2212 H, where i = 2, . . . , n, n is the number of elements in the link. In the equation of the balance of moments (25.5), one unknown \u03b12 is the direction of action of the force F2 from the side of element 2 at the contact point of elements 2 and 3 Mc2. In this case, the solution is found numerically in the simulation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000996_0954410019846713-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000996_0954410019846713-Figure1-1.png",
        "caption": "Figure 1. F-18 high-alpha research vehicle (NASA photo number EC89-0096-149) and assumed variable-span geometry.",
        "texts": [
            " The body-axes aerodynamic forces XA, YA, and ZA can be represented in terms of the wind-axes aerodynamic forces D, C, and L as follows XA \u00bc D cos cos \u00fe C cos sin \u00fe L sin \u00f013\u00de YA \u00bc D sin C cos \u00f014\u00de ZA \u00bc D sin cos \u00fe C sin sin L cos \u00f015\u00de The wind-axes aerodynamic forces are defined as D \u00bc qSCD, C \u00bc qSCC, L \u00bc qSCL \u00f016\u00de where the dynamic pressure is given as follows q \u00bc 1 2 V2 T \u00f017\u00de Assuming that the engine thrust is aligned with the body x-axis, the wind-axes moments are defined similarly as follows l \u00bc qSbCl \u00fe zrefYA \u00f018\u00de m \u00bc qS cCm xref\u00f0L cos \u00feD sin \u00de \u00fe zref\u00f0 L sin \u00feD cos \u00de \u00f019\u00de n \u00bc qSbCn xrefYA \u00f020\u00de The model modification considering variable-span morphing shown in Figure 1 is now introduced. Morphing parameter 2 \u00bd0, 1 is defined as the span extension ratio. In other words, \u00bc 0 when the moving part is completely retracted, and \u00bc 1 when fully extended. The variable-span morphing has a major impact on the external shape of the main wing, and the shape can be characterized by geometrical parameters: planform area, mean aerodynamic chord, and span. Here the morphing factors S, c, and b are defined as follows S\u00f0 \u00de \u00bc S\u00fe cmb S \u00f021\u00de c \u00f0 \u00de \u00bc cb\u00fe cmb cb\u00fe cb \u00f022\u00de b \u00f0 \u00de \u00bc b\u00fe b b \u00f023\u00de Note that they represent the ratio of the morphed geometrical parameter to the nominal (retracted) parameter"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003907_nme.1620380611-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003907_nme.1620380611-Figure2-1.png",
        "caption": "Figure 2. Geometry of a circular helix",
        "texts": [
            " (224 (22b) (22c) (22d) W e ) (22f) Equations (21) and (22) give a set of 12 first-order ordinary linear differential equations which govern the statical behaviour of pre-twisted spatial rods on elastic foundation under nonisothermal conditions including the effects of warping, axial and shear deformations. 3. FIELD EQUATIONS OF THE NON-CIRCULAR HELICOIDAL BAR AND THEIR SOLUTION BY THE COMPLEMENTARY FUNCTIONS METHOD (CFM) The parametric equation of a helix is x = acos4, y = asin&, z = h4 The infinitesimal length element of the helix is defined as ds = (a2 + h2)1\u20192 d+ = c d 4 The relationships between the moving axes and the fixed reference frame are (Figure 2) {t,n,b}T = CB1 {i,j,k}T (254 where - (a/c) sin 4 (a/c) cos 4 - sin 4 (h/c) sin + - (h/c) cos 4 a/c 1038 V. HAKTANIR The curvatures of the circular helix are then given as x = a/cz, z = h/c2 The Frenet formulae for the helix are dt/d$ = (a/c)n, dn/d4 = (h/c)b - (a/c)t, db/d+ = - (h/c)n (27) The curvatures of the non-circular helicoidal bars are not constant along the axes. For example, radius of the helicoid for conical helix is given by (Figure 7) Variation of radius along the axes of hyperboloidal or barrel helix can be found in Reference 12"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000447_jmes_jour_1960_002_027_02-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000447_jmes_jour_1960_002_027_02-Figure4-1.png",
        "caption": "Fig. 4. Complete force field for tube ironing",
        "texts": [
            " Different values for the upper bound will be obtained from different positions of the points a, b, and c, and the best upper bound will be that which gives the lowest values of the forces P, F,, and F2. For any given position of b, it can be shown that the best positions of a and c are at the beginning and end of the die contact length, as shown in Fig. 3. There is no direct method Vol2 No 3 1960 at UNIV NEBRASKA LIBRARIES on June 5, 2016jms.sagepub.comDownloaded from however, of obtaining the best position of point b. In Fig. 4a the force polygons of Fig. 2c are combined to give the complete force field for the upper-bound condition. This field may be constructed from Fig. 3 as shown by the dotted lines in Fig. 4a; alblcldl is a parallelogram. Force fields for several other positions of b (Fig. 3) and bl (Fig. 4a) may be superimposed on Fig. 4a and the best or minimum value of the drawing force P obtained by plotting the locus of point el and drawing the appropriate tangent (Fig. 4b). The mean drawing stress ratio p/2K is quickly obtained from this diagram since p/2K = P/2K. h. In the absence of friction the force field is shown in Fig. 4c. For this special case, an analytical solution to the problem may be found. Any position of bl may be denoted by angles + and 0 (Fig. 4c). If I) = 0, then the upper bound is _ - P sin a cos a 2~ - sin (e+ sin (e+,) For general values of I) and 6 f=F{ 1 sin e sin (e--a)+sin and the best upper bound is _ - P I 2K smcr y - -{y+i-2 cos a} wherey = z/ (h/H). These equations are identical to those obtained by Hill (I) and Green (2) using the same configuration as a velocity J O U R N A L MECHANICAL E N G I N E E R I N G S C I E N C E Vol2 No 3 1960 at UNIV NEBRASKA LIBRARIES on June 5, 2016jms.sagepub.comDownloaded from UPPER-BOUND VALUES FOR THE LOADS ON A RIGID-PLASTIC BODY IN PLANE STRAIN 181 discontinuity pattern",
            " d ( h / H ) . Further minimization shows that the best upper wound is obtained when angles and #* are each equal to ilr, the corresponding optimum value for a total reduction J O U R N A L M E C H A N I C A L E N G I N E E R I N G S C I E N C E 2 ratio of z/(h/H). It is assumed that th is rule applies to the frictional case also. For example, if sheet is reduced in thickness by 36 per cent by drawing through a 10\" die, h/H = 0.64 and d ( h / H ) = 0.80. Hence a reduction of 20 per cent is considered (Fig. 4b) and the resulting optimum value of t,b, 5 2 9 , is used to obtain the complete force field for 36 per cent reduction, as shown in Fig. 56. This equation is derived fiom an equilibrium stress distribution satisfying the yield criterion and the stress boundary conditions in accordance with the principle of maximum plastic work (I). Fig. 7a shows sheet being drawn between taper dies, with supposed dividing planes ab, ac, ad, and ae. Four planes are necessary to obtain a realistic upper bound. The force field for the configuration is shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001779_lra.2020.3007467-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001779_lra.2020.3007467-Figure2-1.png",
        "caption": "Fig. 2. Left: The tendon transmission of an underactuated finger is dependent on pulley and spring parameters. Right: Object geometry can be generalized by evaluating the triangle relationship, T , between the contacts, and offsetting the manipulation frame, M, from the grasp frame, X .",
        "texts": [
            " Given an actuation velocity, a\u0307, and the grasp frame, Xt, at time t, the energy-based propagation model (or system dynamics model) provides a prediction for the next step of the grasp frame pose, Xt+1. This transition is calculated given a tendon transmission constraint, raia\u0307i = rpiq\u0307pi + rdiq\u0307di (3) and the contact triangle constraint, Tt = Tt+1. Thus, we can find the equilibrated joint configuration of the hand, q\u2217 by, q\u2217 = argmin \u2211 i Ei(qi) s.t. (2), (3) (4) where Ei is the potential energy of the ith finger, Ei(qi) = 1 2 (kpq 2 pi + kdq 2 di) (5) Here, rpi, rdi, and rai are the radii of the pulleys on the proximal joint, distal joint, and actuator, respectively, on finger i (Fig. 2). Similarly, \u02d9qpi, \u02d9qdi, and a\u0307i are the rotational velocities about the same joint on the same finger. This energy-based propagation model enables efficient data collection in simulation, and has shown to easily transfer to a physical system [12]. By predefining various contact relationships in T and applying a random actuation input, a\u0307, we observe the grasp frame transition from Xt to Xt+1, thus calculating X\u0307 \u2208 se(3) by taking the element-wise difference. With a 15-dimensional input feature, sn = (Xn, X\u0307n, Tn), and an output feature, a\u0307n, we build the training set, S = {sn}n=1:N , R = {a\u0307n}n=1:N where N denotes training sample size",
            " Algorithm 3 perturb(\u00b7) Input: X\u0307t Output: X\u0307 \u2032 t+1 1: \u03b4x, \u03b4y, \u03b4z \u2190 translationalLimit 2: \u03b4\u03b8R , \u03b4\u03b8P , \u03b4\u03b8Y \u2190 rotationalLimit 3: for i in [x, y, z, \u03b8R, \u03b8P , \u03b8Y ] do 4: X\u0307t+1 \u2190 X\u0307t + rand.uniform(\u2212\u03b4i, \u03b4i) 5: return X\u0307 \u2032 t+1 The proposed control framework was instantiated on a 3- fingered underactuated Yale Openhand Model O. Physical modifications to the readily available open source design include a rounded fingertip and pulleys/bearings within the finger as to reduce friction in the tendon\u2019s transmission. Each finger, composed of two links, is actuated by a single Dynamixel XM-430 motor with return forces supplied by springs at each of the joints (Fig. 2). The learned model in (6) was trained with a dataset of size 300,000 over 50 different contact triangles, T , by evaluating the input-output relationship after random actuation of the energy model in (4). A Random Forest model of tree depth 10 and forest size of 30 was trained, which accounted for joint limits and actuation constraints. Due to the different values in T used for training, the learned model was able to generalize over different object geometries, which is beneficial as it enables adaption to undesired contact scenarios where the relational geometry between the fingertips change, e.g. rolling or slip, as previously presented in [12]. We implemented translational control, i.e. c = (x, y, z), in a simulated environment (Fig. 2) while varying the control horizon and number of optimization iterations as to tune the controller. This test, presented in Fig. 4, tracks the x, y, z position of the manipulation frame over time in an attempt to trace the letters \u2019GRABLAB\u2019. Depicted in different colors, three different-sized objects were used in experimentation, with properties presented in Fig. 5. Each letter was 20mm in height and 10mm in width and was written within the x \u2212 y plane. Letters were comprised of a number of goal points\u2013 squares (start), circles (intermediate), and stars (end)\u2013with 50 waypoints in between each goal",
            " In addition to a purely translational trajectory about the manipulation frame, we test the control approach with other partially constrained trajectories, namely, a purely rotational trajectory c = (\u03b8R, \u03b8P , \u03b8Y ), and a mixed trajectory, c = (z, \u03b8R, \u03b8Y ). This choice of trajectories further underscores the diversity of dimensional combinations which can be inherently accounted for in this framework, after retuning weighting parameters in the cost function and scaling the controlled dimensions to characteristic length. In each of these tests, the hand was initialized with the same hand configuration as in Fig. 2, using Obj. 1. With a horizon of 3 and with 50 optimization iterations, a goal trajectory was formed transitioning M from its current state to a goal configuration. Five trials were executed, resetting the hand after each trial. We record the state of the manipulation frame along the execution trajectory. As presented in Fig. 7, the trajectory ofM was able to successfully follow the desired Authorized licensed use limited to: Auckland University of Technology. Downloaded on July 14,2020 at 12:07:51 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002922_j.mechmachtheory.2021.104382-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002922_j.mechmachtheory.2021.104382-Figure3-1.png",
        "caption": "Fig. 3. Map between acceleration space of the end-effector and driving torque space.",
        "texts": [
            " In rigid dynamics, the inertial properties describe the resistance to acceleration experienced by a force acting on the endeffector. The relationship between acceleration of the end-effector and driving torque can be obtained from the equations of motion. In this section, a parallel manipulator actuated by revolute joints is considered as an example. When the endeffector accelerates with unit translational acceleration in an arbitrary direction, the map between the acceleration space of the end-effector and the driving torque space can be represented as shown in Fig. 3 . The relationship between the driving torque and translational acceleration can be expressed as \u03c4ta = M a t (21) where \u03c4ta represents the driving torque caused by the translational acceleration of the end-effector, a t is the unit translational acceleration, and M is the mass matrix of the manipulator. The singular value decomposition of M can be derived as M = U V T (22) where U and V are 2 \u00d7 2 orthogonal matrices, and is a 2 \u00d7 2 diagonal matrix given by = [ \u03c31 0 0 \u03c32 ] , \u03c31 \u2265 \u03c32 (23) The scalars \u03c31 and \u03c32 are singular values of M , and they represent the lengths of the major and minor axes of the ellipse in the driving torque space (shown in Fig. 3 ), respectively. As shown in Fig. 4 , ellipses I and II are different responses in the driving torque space corresponding to the unit translational acceleration. In addition, the lengths of the major and minor axes are the same for the two ellipses. The dynamic performance of these two statuses shows no difference with respect to the dynamic conditioning index and dynamic manipulability. However, the discrepancy between the maximum driving torques ( \u03c41 and \u03c42 in Fig. 4 ) in ellipse I is larger than that in ellipse II"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000885_1.4042867-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000885_1.4042867-Figure1-1.png",
        "caption": "Fig. 1 DMLS SS GP1 sample manufactured: (a) sample design for DMLS manufacturing [14], (b) support structure of sample, (c) layout of as-built DMLS manufactured SS GP1 samples on build plate with numbering [14], and (d ) experimental setup for tension and fatigue testing [8]",
        "texts": [
            " In order to assess the monotonic response, DMLS SS GP1 samples of each build orientation were subject to room temperature strain-controlled monotonic tension tests [13] at a strain rate of 0.001 mm/mm/s, and a sampling rate of 25 Hz. All tension tests were performed using the servohydraulic MTS LandMark 793 test system, and a clip on extensometer, MTS 647.11E25, was used to record strain measurements up to a strain value of 0.2 mm/mm, after which measurements were recorded in displacement control until sample fracture [14]. The experimental setup for tension and fatigue testing is as shown in Fig. 1(d ). After 0.2 mm/mm strain, displacement values were used to calculate the average strain to fracture, as done in a previous study on DMLS Inconel 718 [15]. Tension tests were interrupted during testing and therefore were completed in two tests. Stress\u2013strain curves from both tests were combined to yield the complete monotonic response. Samples of each build orientation were subject to room temperature strain-controlled, completely reversible (R\u03b5=\u22121) LCF tests at a strain rate of 0.001 mm/mm/s, strain range of, \u0394\u03b5 1",
            " The samples were designed with an inner gauge diameter of 6.35 mm and a gauge length of 25.4 mm, as suggested by ASTM standards [13,19], and at varying build orientations across the xy build platform [8]. The sample support was set to a height of 5 mm. The samples that were intended to be manufactured along the x-axis/(100) direction and y-axis/(010) directions, respectively, were subject to a \u22125 deg offset for these orientations as suggested by the manufacturer. The layout of the samples across the build plate is depicted in Fig. 1(c), along with the sample numbers specific to each orientation. The samples were not stress-relieved/heattreated after manufacturing, in order to investigate the as-built mechanical performance of this material. Delaminations are often observed in additively manufactured components and are suggested to be attributed to stress relief, due to a buildup of high residual stresses during the DMLS manufacturing process. The presence of delaminations can be distinctly observed from Fig. 2(a) in the outer shank region of the samples\u2019 support structure, which may be attributed to stress relief or to the sharp change in geometry of the sample outer shank diameter from 12",
            " Figures 2(a) and 2(b) show an image of a sample with the boxed support structure attached. After removal of the support structure, as indicated in Fig. 2(c), support structure remnants and grayish unmelted powder can be observed. These remnants were removed through filing of the samples, using a flat smooth file, in the location where the boxed support structure was originally present, as depicted in Fig. 2(d ). The as-built DMLS SS GP1 samples manufactured along varying build orientations in the xy plane (see Fig. 1(c) for reference frame) were observed to be slightly warped in the outer gauge/shank section of the samples. The angle of warping for each sample was measured using a protractor and found to be minimal, 1 deg to 2 deg. The boxed support structure used for DMLS manufacturing of these samples can be attributed to minimizing the angle of warping. Warping of additively manufactured samples is commonly observed for samples manufactured in the xy build plane and has been attributed to the thermal gradients experienced by these samples [15], during the layer-by-layer melting and cooling process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001776_tie.2020.3005108-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001776_tie.2020.3005108-Figure13-1.png",
        "caption": "Fig. 13. Field-circuit coupling simulation model.",
        "texts": [
            " The maximum fundamental component of non-modular and modular VFRM is 12s10r and 12s14r, respectively, as shown in Fig. 12 (b). Additionally, the 10- and 14-rotor-pole nonmodular VFRM do not exhibit third-order harmonic, which can be illustrated by their winding factor of third-order harmonic is zero. Besides, the modular VFRM have thirdorder harmonic due to the flux gap changes the coil pitch factor, which can be explained by the formula (6). A three phases full wave rectifier is adopted in field-circuit coupling simulation model, as shown in Fig. 13, where C is a voltage regulated capacitance. All the VFRM rotates at speed 500 r/min. Fig. 14 (a) shows the output voltage versus output current at same DC-field current for non-modular and modular VFRM, and a varying resistance load is adopted. It can be seen from Fig.14 (a) that the output voltage of modular 12s10r VFRM at 0.2A decreases about 28.3% compared with nonmodular VFRM. The modular 12s14r VFRM increases about 40%, which is accorded with the foregoing analysis, i.e., the fundamental component of flux linkage and back-EMF are increased in 12s14r modular VFRM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001184_j.ijfatigue.2019.105281-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001184_j.ijfatigue.2019.105281-Figure14-1.png",
        "caption": "Fig. 14. 3D elliptical crack in an infinite body under the uniform shear stress \u03c4 [35,36].",
        "texts": [
            " The difference in the test results is probably due to the difference in the test materials. In this section, the initiation behaviour of the subsurface RCF crack was studied based on the fracture mechanics approach, considering the actual morphology of the macroscopic defect. Because the growth of the mode I crack is suppressed by the compression stress under wheel\u2013rail contact, the subsurface RCF crack mainly propagates in a mixed mode (II\u2013III) [19,20]. The macroscopic defect in the wheel rim can be assumed to be a 3D elliptical crack. Fig. 14 shows a 3D elliptical crack in an infinite body under the uniform shear stress \u03c4. The tip of the major axis of the elliptical crack (point A) is subjected to pure mode II loading, while the tip of the minor axis of the elliptical crack (point B) is subjected to pure mode III loading. The parameters KII at point A and KIII at point B can be expressed as [35]: = \u2212 + \u2032 \u2032K k \u03c4 k v E k vk K k \u03c0k S ( ) ( ) ( )II 2 2 2 34 (2) = \u2212 \u2212 + \u2032 \u2032K v k \u03c4 k v E k vk K k \u03c0k S(1 ) ( ) ( ) ( )III 2 2 2 4 (3) where, \u222b= \u2212K k k \u03d5 d\u03d5( ) (1 sin ) \u03c0 0 /2 2 2 1/2 \u222b= \u2212 E k d\u03d5 k \u03d5 ( ) (1 sin ) \u03c0 0 /2 2 2 1/2 a is the semi-major axis, b is the semi-minor axis, v is the Poisson\u2019s ratio, k2= 1\u2212b2/a2, k\u20192= b2/a2, and S= \u03c0ab"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003680_1999-01-0404-Figure21-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003680_1999-01-0404-Figure21-1.png",
        "caption": "Figure 21. Visteon Environmental Dynamometer facilty",
        "texts": [
            " Environmental Testing \u2013 As part of the design validation process, the EHPAS system and entire vehicle is stressed and tested functionally under multiple environmental conditions. The ECM is given a classification in accordance with the area on the vehicle in which it will be packaged. These classifications define the temperature limits, water ingress severity, and severity of other contaminants that the system could be exposed. Many of these tests have been conducted in Visteon environmental Dynamometer Facility Fig. 21, that has capabilities to provide different driving and environmental conditions. The environmental stresses include thermal shock to test mechanical stress robustness (transitioning from -40\u00b0C soak temperature cycles to 125\u00b0C temperature cycles), high and low temperature operation testing (exercising the EHPAS system at the extreme temperature limits), powered thermal cycle (exercising the electronic module while cycling between the temperature limits), vibration testing, humidity testing (exercising the module while exposed to high humidity conditions), water submersion 10 testing, salt spray exposure, and dust exposure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000721_j.ymssp.2016.09.033-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000721_j.ymssp.2016.09.033-Figure3-1.png",
        "caption": "Fig. 3. The size and location of the defect in the tooth.",
        "texts": [
            " 2, the pinion gear and wheel gear under study have five flexible teeth, while the rest are rigid components that will not be deformed in the calculation. The subsurface defect is located at one of the flexible teeth of pinion gear, whose stress and strain results are of the main concern. In order to locate the subsurface crack, a block with 0.5 mm\u00d70.5 mm\u00d70.05 mm is removed from the face of the defective tooth near the pitch line, and one crack in the shape of the coin is located on this block, see Fig. 3. The crack is set on the subsurface, and the original sizes are 2a=0.05 mm, 2b=0.09 mm, h=0.01 mm and d=0.04 mm. The original sizes and depths of the subsurface crack are defined after [10,13]. To avoid the geometrical incompatibilities of the deflected tooth, the defective tooth is divided into eight regular parts. After that, the block (where the subsurface crack locates) and the rest part of this tooth are grouped into one multibody part using \u201cform new part\u201d tool in ANSYS Workbench, which enables the use of shared topology among the bodies to obtain high quality mesh",
            "020 mm are more appropriate. For more precise results, 0.015 mm is selected as the optimal element size. With the element size of 0.015 mm under the condition of T=100 N m and f=0.06, the calculated cycle numbers to each state are listed in Table 4. The tooth contact pressure maps relative to the fault location for state I\u2013III is shown in Fig. 11 at t=0.19\u00d710\u22123 s and 0.27\u00d710\u22123 s respectively. Details are given in the Supplementary material. Herein, state I is the initial state of the gears as illustrated in Fig. 3. The crack size and depth for each state are also given with their appearances in Fig. 12. After the third iterative in state III, the model is modified for the third time and a cavity is formed on the contact surface, which is state IV. Under the conditions of f=0.04 and 0.06, the external torque T=100 N m and the pinion speed of 5000 rpm, the time-varying global maximum values of accumulated plastic strain distributed over the defective tooth in one mesh cycle for the first three state (state I, II, III in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000934_tmag.2019.2900527-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000934_tmag.2019.2900527-Figure5-1.png",
        "caption": "Fig. 5. Basic configuration of a three-phase-laminated-core variable inductor.",
        "texts": [
            " VARIABLE INDUCTOR INCORPORATING PLAY MODEL In this section, first, a basic configuration and an operating principle of a three-phase-laminated-core variable inductor are described. Next, a method for deriving the RNA model of the variable inductor incorporating the play model is explained. Three-Phase-Laminated-Core Variable Inductor Variable inductors consist of only a magnetic core and primary dc and secondary ac windings [12]. It can control the winding inductance quickly and continuously by the magnetic saturation effect. Thus, it can be used as a reactive power compensator by combining with fixed power capacitors in electric power systems [13]. Fig. 5 illustrates a basic configuration of a three-phaselaminated-core variable inductor [11], which consists of six legs, ring yokes, and primary dc and secondary ac windings. The secondary ac windings Nu , Nv , and Nw are coiled around each couple of legs on the same line, respectively. These windings are connected to delta and to the three-phase ac source. On the other hand, the primary dc windings N1 are coiled around the outside ring yoke so that the dc flux circulates in the yoke, and connected to the dc source Vdc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003029_j.ymssp.2021.108116-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003029_j.ymssp.2021.108116-Figure1-1.png",
        "caption": "Fig. 1. Reference CCHGP design used in this paper. (a) the transverse gear profile inside the casing, showing the angular convention and the definition of the casing start/end angle; (b) profile of a single tooth for the designed circular-arc/involute compound gear profile. The profile between Points B and C is involute; (c) a 3-D CAD model of the sample gear pump design; (d) axial balancing mechanism in the reference design; (e) components of the reference unit, shown in an exploded view.",
        "texts": [
            " In addition, the dynamic characteristics will be compared and discussed in different hydrostatic balancing design scenarios, shedding light on the appropriate balancing strategy of CCHGP as a highly promising hydraulic pump design. A continuous-contact helical gear pump (CCHGP) is a category of EGPs that utilize a fully conjugate gear profile and helical gears to reduce, or eliminate, the kinematic flow ripple. The CCHGP concept can be implemented with various morphologies of gear profile, as summarized in [20]. The gear profile used for the reference design discussed in this paper is the circular-arc/involute compound design, as shown in Fig. 1b. In this profile, the tip and the root of the tooth are circular arcs with the same radius of curvature, and a portion of the involute profile is used to connect two circular arcs. The involute curve is defined by the pressure angle on the transverse plane \u03b1t. To achieve the continuity in the curvature at the connection points (B and C, in Fig. 1b), the relation between the transverse pressure angle \u03b1t and the addendum and dedendum values (i.e. outer radius and root radius) can be determined, with a mathematical derivation that can be found in [20]. Due to the nature of the continuous-contact gear profile, the transverse profile contact ratio of this family of gear profile is 0.5, at the designed center distance. For gears to drive themselves without an external drive, the overall contact ratio needs to be greater than 1.0. For this reason, a CCHGP requires helical gears, with a helical contact ratio no smaller than 0.5. As deeply discussed in [20], a helix contact ratio of 1.0 will give the cancellation of the kinematic flow ripple. The total contact ratio is the sum of transverse contact ratio and helical contact ratio, and therefore 1.5 for the reference pump discussed in this paper. The fully conjugate nature of the gear profile eliminates the residual volume trapped between the gears during the meshing process (as shown in Fig. 1a). For this reason, the operation of this type of EGP does not rely on lateral relief grooves, which are required in traditional involute EGPs, to avoid fluid volumes being trapped between points of contact. As shown in Fig. 1e, the reference CCHGP unit consists of two rotors, and two pairs of bearing blocks, one pair on each side of the rotors. The axial balancing of the rotors for the reference pump design is implemented by putting two balancing pistons on one end of each rotor, and the outer side of the piston is connected to the delivery port via a microchannel inside the cover plate (Fig. 1d and e). The proposed model in this paper is capable of modeling the dynamic behavior of the supporting mechanism of this type. There are other axial balancing solutions that have been seen on commercial products (for example, asymmetric design of relief grooves on the inner side of bearing blocks, used by SILENCE PLUS by Bosch Rexroth [28]), but they are not in the scope of the current work. X. Zhao and A. Vacca Mechanical Systems and Signal Processing 163 (2022) 108116 The pump taken as reference in the present study has design parameters summarized in Table 1. Some of the design parameters and a CAD model of the reference design are shown in Fig. 1c. For the reference design is the value of the pitch circle helix angle, which is 21.44\u25e6 so that the helical contact ratio is equal to 1.0. As this type of gear pump does not need any relief grooves, therefore the lateral bushings of the reference design (shown in golden color in Fig. 1c) have a flat surface on the side facing the gears. The coordinate system used in this paper, without further mentioning, is shown in Fig. 1a: the center of the coordinate is located at the center of the driver gear\u2019s cylinder, with the same distance to the top and bottom plane of the gear. The z-axis is the axial direction of the gear, while x- and y-axis are on the transverse plane of the gear. x-axis is parallel to the line connecting the centers of two meshing gears, to which y-axis is perpendicular. For helical gears with a positive helix angle, the gear profile on the top surface of the top has a leading phase angle to the bottom surface",
            " For the force components on the transverse plane, due to non-uniform application area distribution along the axial direction, which leads to a distribution of forces, a moment with respect to x- and y-axis can be generated, (Fig. 4a and b); on the other hand, because of the helical gear structure, the fluids surrounding the gear tends to push the leading flank of one tooth to one axial direction and push the trailing flank of the adjacent tooth to the opposite direction (Fig. 4c), which creates a moment and a net axial loading. The axial rotor surfaces (Fig. 5a) in helical gear pumps are in direct contact with the axial lubrication gaps. The latter is the lubrication gap between the gears and the bearing blocks (Fig. 1e), typically with a thickness on the order of 10 \u03bcm. When the pump is operating and the gears are spinning at high speed, the thin layer of fluid film in these gaps provides the bearing function to support the load, as well as the sealing function to prevent the working fluids from leaking to the outside. For CCHGPs, the fluid-pressure distribution in the top and bottom gaps is different, due to two reasons: (1) helical gears usually have different profile angles for their top and bottom surfaces, resulting in a phase-delay between the top and bottom gaps; (2) due to the presence of sealing surface in the meshing zone of CCHGPs, the same tooth-space could have extremely high pressure-difference between its top end and bottom end",
            " For high-pressure hydraulic applications, the forces on the gear body are so high that inertia effects can be neglected, which is to say the gear body is assumed to be statically balanced at every instant. The force balance in z-direction is written as: Fz,balance +Fz,p1 +Fz,c1 +Fz,p2,top(vz)+Fz,p2,bot(vz) = 0 (52) where Fz,balance is the balancing force given by the balancing mechanism, if any. For CCHGPs, an axial balancing design is typically needed (as in the reference design considered in this paper, Fig. 1), because the load given by the fluid pressure and the gear contact is highly asymmetric. To study the gear dynamics and aid the axial balancing design, two scenarios are discussed: axial static scenario: all the axial loads on the gears are perfectly balanced by the supporting force Fz,balance, so that at every moment the gear is statically balanced right in the middle, with equal lubrication film thickness on top and bottom side (i.e. vz = \u0394z = 0,htop = hbottom). This is an ideal balancing condition to pursue, and this scenario can be used to determine the ideal amount of Fz,balance, as discussed later in Section 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000245_j.jmatprotec.2019.116515-Figure19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000245_j.jmatprotec.2019.116515-Figure19-1.png",
        "caption": "Fig. 19. Equivalent stress distribution obtained from of the 3rd pass spinning simulation (top view): a) conventional spinning; b) SSFC spinning; c) DSFC spinning; d) comparison of equivalent stress for the three processes.",
        "texts": [
            " Based on the results of finite element calculation, this section focuses on the mechanism analysis of the thickness uniformity improvement in the multi-pass spinning. Fig. 18 shows the wall thickness values of the forming region where the material has been fitted to the mandrel from the 1st, 3rd, 5th, 7th pass simulations during the 8-pass spinning. It can be seen that the maximal difference of thickness between the conventional and DSFC processes distributes within the range from 45\u00b0 to 70\u00b0, namely from the 3rd to 6th pass. The contours of equivalent stress in the third pass simulation for the three spinning processes were shown in Fig. 19a-c, in which the equivalent stress in the gray area is less than 125 MPa. Comparing between the three contours in Fig. 19a-c, it is found that the flangeconstrained spinning processes, especially the DSFC spinning, can make more material at the flange area to produce plastic deformation. Fig. 19d displays the equivalent stress along a section from the simulations of the third pass spinning. In the flange-constrained spinning processes, the equivalent stress distribution was more uniform, especially in the DSFC process. Compared with the conventional spinning, the equivalent stress is smaller in the forming region and larger in the flange region during the DSFC spinning, which makes the material flow from the flange to the forming region more easily. For the SSFC spinning, the similar conclusion is also obtained"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000087_1750-3841.14752-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000087_1750-3841.14752-Figure6-1.png",
        "caption": "Figure 6\u2013(A) Response surface and (B) contours for extraction time and temperature.",
        "texts": [
            "1202 show that the lack of fit was insignificant (P > 0.05), hence, the model is stable and reliable. Meanwhile, the coefficient of determination (R2 = 0.9778), adjusted coefficient of determination (R2 adj = 0.9494), and coefficient of variance (CV = 4.27%) clearly confirmed that the model was highly significant. Analysis of the response surface and contour plots. Three-dimensional response and contour plots were constructed to visualize the relationship between the variables and extraction efficiency (E). The results are shown in Figure 6 to 8. The slope of the surface and sparsity of the contours were used to determine the relationship between any two variables and the optimum conditions for the maximum extraction efficiency. Figure 6A shows the effects of the temperature and time on the extraction efficiency, which reached a maximum of 92.11%. The contour plot in Figure 6B confirms that the mutual interaction between temperature and time was insignificant (P = 0.7106 > 0.05). As shown in Figures 7A and 8A, the interaction effects of AC and BC had the maximum impact on the extraction efficiency. The circular contour plots in Figure 7B indicate that the mutual interaction between the extraction temperature (A) and power (C) was quite significant, which supports the results given in Table 2 (P = 0.0102 > 0.05). In Figure 8B, the 2-D elliptical contour plot reveals that the mutual interaction between the extraction time (B) and power (C) was significant (P = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000791_j.ifacol.2016.10.194-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000791_j.ifacol.2016.10.194-Figure2-1.png",
        "caption": "Fig. 2. Acrobot model",
        "texts": [],
        "surrounding_texts": [
            "IFAC-PapersOnLine 49-18 (2016) 374\u2013379\nScienceDirect\nAvailable online at www.sciencedirect.com\n2405-8963 \u00a9 2016, IFAC (International Federation of Automatic Control) Hosting by Elsevier Ltd. All rights reserved. Peer review under responsibility of International Federation of Automatic Control. 10.1016/j.ifacol.2016.10.194\n\u00a9 2016, IFAC (International Federation of Automatic Control) Hosting by Elsevier Ltd. All rights reserved.\nSwing up and stabilization of the Acrobot via nonlinear optimal control based on\nstable manifold method\nTakamasa Horibe \u2217 Noboru Sakamoto \u2217\u2217\n\u2217 Department of Aerospace Engineering, Graduate School of Engineering, Nagoya University, Furo-cho, Chikusa-ku, Nagoya,\nJapan, (e-mail: horibe.takamasa@j.mbox.nagoya-u.ac.jp) \u2217\u2217 Department of Mechatronics, Faculty of Science and Engineering, Nanzan University, Yamazato-cho 18, Showa-ku, Nagoya, Japan,\n(e-mail: noboru.sakamoto@nanzan-u.ac.jp)\nAbstract: This paper considers the problem of swing up and stabilization for the Acrobot. It is shown that stable manifold method which has been proposed for computing nonlinear optimal control is capable of designing feedback controllers for this problem. An optimal stabilization controller is obtained as a single feedback law by numerically solving a Hamilton-Jacobi equation by the stable manifold method. It is shown that unlike existing methods for Acrobot swing up such as partial feedback linearization, the resultant control is mechanically indigenous in the sense that it uses reactions of arms effectively and, as a consequence, control input is kept low. A number of simulations verify the effectiveness and robustness of the controller.\nKeywords: Acrobot, Nonlinear optimal control, Hamilton-Jacobi equation, stable manifold method\n1. INTRODUCTION\nUnderactuated systems are mechanical systems which have fewer control inputs than degrees of freedom. Control of underactuated systems is currently an active topic for many researchers due to wide application range in Robotics or aerospace field (Liu and Yu, 2013; Xin and Liu, 2013). The Acrobot is a 2-dimensional underactuated mechanical system often used as a benchmark problem for testing nonlinear control methods. It consists of two links and an actuator is installed only at the second joint. A common control objective of the Acrobot is to swing it up from the downward position to the unstable upright position and to stabilize it vertically. This is a challenging task because of the movement in a large range of nonlinearity.\nGenerally, the swing up control is divided into two phases, first, the swing up phase in which nonlinearity is dominant, and then, stabilization phase which estabishes autonomus stability in a neighborhood of the origin. There are some other effective ways to design swing up controllers such as using partial feedback linearization (Spong, 1995), energy feedback (Xin and Yamasaki, 2012; Xin and Kaneda, 2007), trajectory tracking (Zhang et al., 2013), Lyapunov based control (Zergeroglu et al., 1998) and intelligent control (Brown and Passino, 1997). However, switching controller has no guarantee of stability in the vicinity of the boundary. Researchers in (Davison and Bortoff, 1997; Xin and Kaneda, 2001) propose methods to enlarge the region of attraction (RoA) of linear controllers, stabilization of the Acrobot by linear control is inherently difficult. Backstepping approachOlfati-Saber (2000) should be\nmentioned as a single feedback control method under some assumptions that are difficult to satisfy generally.\nIn this paper, we show that it is possible to design a single (without switching) optimal feedback controller for swing up and stabilization of the Acrobot using the stable manifold method (Sakamoto and van der Schaft, 2008; Sakamoto, 2013). The method has been developed for numerically computing the derivative of solution for Hamilton-Jacobi equations (HJEs). When it is applied for the Acrobot swing up problem, it directly enlarges the RoA for stabilization so that the downward position is included in RoA. For a survey and other solution methods for HJEs, we refer to Aguilar and Krener (2014); Aliyu (2011); Beeler et al. (2000); Lukes (1969); Navasca and Krener (2007).\nThe organization of the paper is as follows. The Acrobot model is introduced in \u00a7 2. \u00a7 3 summarizes the theory of the stable manifold method for HJEs. Controller design is precisely explained in \u00a7 5 and simulation results are shown in \u00a7 6.\n2. MODELING AND ANALYSIS OF THE ACROBOT\nIn this section, we derive a nonlinear model of the Acrobot. Figs. 1, 2 show the Acrobot and its schematic model. The control torque is applied only to the second joint from an actuator through a pulley and a timing belt.\nFor the ith link (i = 1, 2), qi is the angle, mi is the mass, li is the length, lci is the distance from ith joint to the center of mass (COM), Ji is the inertia around the center of mass, and let g be a gravitational acceleration (9.801\n10th IFAC Symposium on Nonlinear Control Systems August 23-25, 2016. Monterey, California, USA\nCopyright \u00a9 2016 IFAC 380\nSwing up and stabilization of the Acrobot via nonlinear optimal control based on\nstable manifold method\nTakamasa Horibe \u2217 Noboru Sakamoto \u2217\u2217\n\u2217 Department of Aerospace Engineering, Graduate School of Engineering, Nagoya University, Furo-cho, Chik sa-ku, Nag ya,\nJapan, (e-mail: horibe.takamasa@j.mbox.nagoy u.ac.jp) \u2217\u2217 Department of Mec atronics, Faculty of Science and Engineering, Nanz n University, Y mazato-cho 18, Showa-ku, Nagoya, Japa ,\n(e-mail: noboru.sakam to@nanzan u.ac.jp)\nAbstract: This paper considers the problem of swing up and stabilization for the Acrobot. It is shown that stable manif ld method which has been proposed for computing nonlinear optimal control is cap of designing fee back controll rs for this pr blem. An optimal st bilization ler is obtained as a single f law by numerically solving a Hamilton-Jacobi equ by he stable m ifold method. It is shown that nlike existing methods for Acrobot swing up such as partial feedback linearizat on, the resulta t control is mechanically indigenous i the ense that i uses reactions of arms effectively and, as a consequen e, control put is kept low. A number of simulations verify the ness and robustn ss of the controller.\nKeywords: Acrobot, Nonlinear optimal control, Hamilton-Jacobi equation, stable manifold m th d\n1. INTRODUCTION\nUnderactuated systems are mechanical systems which have fewer control inputs than degrees of freedom. Control o underac uated systems is currently an active topic for many research rs due to wide app ic tion range in Robotics o aerospace field (Liu and Yu, 2013; Xin and Liu, 2013). The Acrobot is a 2-dimensional underactuate mechanical system often used as a benchmark problem for testing nonlinear con rol m thods. It consists of two links and an actuato is installed only a the second joint. A common ontr l objective of the Acrobot i t swing i up fro the downward position o the unstable upright positi n and to stabilize it vertically. This i challen ing ta k because of the movement in a large range of no linearity.\nGenerally, the swing up control is divided into two phases, first, the swing up phase in which nonlinearity is d minant and then, stabilization phase w ich estabishes autonomus stability in neighb rhood of the origin. There are some o her effective ways to design swing up cont ollers such as using partial feedback linearization (Sp g, 1995), energy feedb ck (Xin and Yamasaki, 2012; Xin and Kaneda, 2007), traje tory tr cking (Zh ng et al., 2013), Lyapunov based control (Ze geroglu et al., 1998) and intelligent control (Brown and Passino, 1997). However, switchi g ler has no guarantee of stability in the vicinity of the b undary. Researchers in (Davison and Bortoff, 1997; Xin and Kaneda, 2001) propose methods to enlarge the region of attraction (RoA) f linear controllers, stabilization of the Acrobot by linear control is inheren ly diffi cult. Backstepping approachOlfati-Saber (2000) shoul be\nmentioned as a single feedback control method under some assumptions that ar difficult to satisfy generally.\nIn this paper, we show that it is possible to design a single (without switching) optimal feedback controller for swing up and tabilization of the Acrobot using the stable manifold me hod (Sakamoto and van der Schaft, 2008; Sakam to, 2013). The method has been developed for numerically computing the derivative of solution for Hamilton-Jacobi equatio s (HJEs). When it is applied the Acrobot swing up problem, it dir ctly enlarges the RoA for stabilization so that the downward position is included in RoA. For a survey and other solution method for HJEs, we refer to Aguilar and Krener (2014); Aliyu (2011); Beel r t al. (2000); Lukes (1969); Navasca and Krener (2007).\nThe organization of the paper is as follows. The Acrobot model is introduced in \u00a7 2. \u00a7 3 ummarizes the theory f the stable manifold method for HJEs. Controll r d sign is pr cisely explained in \u00a7 5 and simulation results are hown in \u00a7 6.\n2. MODELING AND ANALYSIS OF THE ACROBOT\nIn this section, we derive a nonlinear model of the Acrobot. Figs. 1, 2 show the Acrobot and its schematic model. The control torque is applied nly to he second joint from an actuator through a ulley and a iming belt.\nFor the ith link (i = 1, 2), qi is the angle, mi is the mass, li is the length, lci is the distance from ith joint to the center of mass (COM), Ji is the inertia around the center of mass, and let g be a gravitational acceleration (9.801\n10th IFAC Symposium on Nonlinear Control Systems August 23-25, 2016. Monterey, California, USA\nCopyright \u00a9 2016 IFAC 380\nS ing up and stabilization of the crobot via nonlinear opti al control based on\nstable anifold ethod\nTakamasa Horibe \u2217 Noboru Sakamoto \u2217\u2217\n\u2217 Department of Aerospace Engineering, Graduate School of Engineering, Nagoya University, Furo-cho, Chikusa-ku, Nagoya,\nJapan, (e-mail: horibe.takamasa@j.mbox.nagoya-u.ac.jp) \u2217\u2217 Department of Mechatronics, Faculty of Science and Engineering, Nanzan University, Yamazato-cho 18, Showa-ku, Nagoya, Japan,\n(e-mail: noboru.sakamoto@nanzan-u.ac.jp)\nAbstract: This paper considers the problem of swing up and stabilization for the Acrobot. It is shown that stable manifold method which has been proposed for computing nonlinear optimal control is capable of designing feedback controllers for this problem. An optimal stabilization controller is obtained as a single feedback law by numerically solving a Hamilton-Jacobi equation by the stable manifold method. It is shown that unlike existing methods for Acrobot swing up such as partial feedback linearization, the resultant control is mechanically indigenous in the sense that it uses reactions of arms effectively and, as a consequence, control input is kept low. A number of simulations verify the effectiveness and robustness of the controller.\nKeywords: Acrobot, Nonlinear optimal control, Hamilton-Jacobi equation, stable manifold method\n1. INTRODUCTION\nUnderactuated systems are mechanical systems which have fewer control inputs than degrees of freedom. Control of underactuated systems is currently an active topic for many researchers due to wide application range in Robotics or aerospace field (Liu and Yu, 2013; Xin and Liu, 2013). The Acrobot is a 2-dimensional underactuated mechanical system often used as a benchmark problem for testing nonlinear control methods. It consists of two links and an actuator is installed only at the second joint. A common control objective of the Acrobot is to swing it up from the downward position to the unstable upright position and to stabilize it vertically. This is a challenging task because of the movement in a large range of nonlinearity.\nGenerally, the swing up control is divided into two phases, first, the swing up phase in which nonlinearity is dominant, and then, stabilization phase which estabishes autonomus stability in a neighborhood of the origin. There are some other effective ways to design swing up controllers such as using partial feedback linearization (Spong, 1995), energy feedback (Xin and Yamasaki, 2012; Xin and Kaneda, 2007), trajectory tracking (Zhang et al., 2013), Lyapunov based control (Zergeroglu et al., 1998) and intelligent control (Brown and Passino, 1997). However, switching controller has no guarantee of stability in the vicinity of the boundary. Researchers in (Davison and Bortoff, 1997; Xin and Kaneda, 2001) propose methods to enlarge the region of attraction (RoA) of linear controllers, stabilization of the Acrobot by linear control is inherently difficult. Backstepping approachOlfati-Saber (2000) should be\nmentioned as a single feedback control method under some assumptions that are difficult to satisfy generally.\nIn this paper, we show that it is possible to design a single (without switching) optimal feedback controller for swing up and stabilization of the Acrobot using the stable manifold method (Sakamoto and van der Schaft, 2008; Sakamoto, 2013). The method has been developed for numerically computing the derivative of solution for Hamilton-Jacobi equations (HJEs). When it is applied for the Acrobot swing up problem, it directly enlarges the RoA for stabilization so that the downward position is included in RoA. For a survey and other solution methods for HJEs, we refer to Aguilar and Krener (2014); Aliyu (2011); Beeler et al. (2000); Lukes (1969); Navasca and Krener (2007).\nThe organization of the paper is as follows. The Acrobot model is introduced in \u00a7 2. \u00a7 3 summarizes the theory of the stable manifold method for HJEs. Controller design is precisely explained in \u00a7 5 and simulation results are shown in \u00a7 6.\n2. MODELING AND ANALYSIS OF THE ACROBOT\nIn this section, we derive a nonlinear model of the Acrobot. Figs. 1, 2 show the Acrobot and its schematic model. The control torque is applied only to the second joint from an actuator through a pulley and a timing belt.\nFor the ith link (i = 1, 2), qi is the angle, mi is the mass, li is the length, lci is the distance from ith joint to the center of mass (COM), Ji is the inertia around the center of mass, and let g be a gravitational acceleration (9.801\n10th IFAC Symposium on Nonlinear Control Systems August 23-25, 2016. Monterey, California, USA\nCopyright 6 AC 380\ni ili i f i li i l l\nl if l\ns ri e \u2217 r t \u2217\u2217\n\u2217 epart e t of erospace gi eeri g, rad ate chool of gi eeri g, agoya iversity, ro-cho, hik sa-k , agoya, Japa , (e- ail: horibe.taka asa j. box. agoya- .ac.jp)\n\u2217\u2217 epart e t of echatro ics, ac lty of cie ce a d gi eeri g, a za iversity, a azato-cho 18, ho a-k , agoya, Japa ,\n(e- ail: obor .saka oto a za - .ac.jp)\nstr ct: is a er co si ers t e ro le of s i g a sta ilizatio for t e cro ot. It is s o t at sta le a ifol et o ic as ee ro ose for co ti g o li ear o ti al co trol is ca a le of esig i g fee ack co trollers for t is ro le . o ti al sta ilizatio co troller is o tai e as a si gle fee ack la y erically solvi g a a ilto -Jaco i eq atio y t e sta le a ifol et o . It is s o t at like existi g et o s for cro ot s i g s c as artial fee ack li earizatio , t e res lta t co trol is ec a ically i ige o s i t e se se t at it ses reactio s of ar s e ectively a , as a co seq e ce, co trol i t is ke t lo .\ner of si latio s verify t e e ective ess a ro st ess of t e co troller.\ney ords: cro ot, o li ear o ti al co trol, a ilto -Jaco i eq atio , sta le a ifol et o\n1. I I\neract ate syste s are ec a ical syste s ic ave fe er co trol i ts t a egrees of free o . o - trol of eract ate syste s is c rre tly a active to ic for a y researc ers e to i e a licatio ra ge i o otics or aeros ace el ( i a , 2013; i a i , 2013). e cro ot is a 2- i e sio al eract ate ec a ical syste ofte se as a e c ark ro le for testi g o li ear co trol et o s. It co sists of t o li ks a a act ator is i stalle o ly at t e seco joi t. co o co trol o jective of t e cro ot is to s i g it fro t e o ar ositio to t e sta le rig t ositio a to sta ilize it vertically. is is a c alle gi g task eca se of t e ove e t i a large ra ge of o li earity.\ne erally, t e s i g co trol is ivi e i to t o ases, rst, t e s i g ase i ic o li earity is o i a t, a t e , sta ilizatio ase ic esta is es a to o s sta ility i a eig or oo of t e origi . ere are so e ot er e ective ays to esig s i g co trollers s c as si g artial fee ack li earizatio ( o g, 1995), e - ergy fee ack ( i a a asaki, 2012; i a a e a, 2007), trajectory tracki g ( a g et al., 2013), ya ov ase co trol ( ergerogl et al., 1998) a i tellige t co trol ( ro a assi o, 1997). o ever, s itc i g co troller as o g ara tee of sta ility i t e vici ity of t e o ary. esearc ers i ( aviso a orto , 1997; i a a e a, 2001) ro ose et o s to e large t e regio of attractio ( o ) of li ear co trollers, sta ilizatio of t e cro ot y li ear co trol is i ere tly i - c lt. ackste i g a roac lfati- a er (2000) s o l e\ne tio e as a si gle fee ack co trol et o er so e ass tio s t at are i c lt to satisfy ge erally.\nI t is a er, e s o t at it is ossi le to esig a si gle ( it o t s itc i g) o ti al fee ack co troller for s i g a sta ilizatio of t e cro ot si g t e sta le a ifol et o ( aka oto a va er c aft, 2008; aka oto, 2013). e et o as ee evelo e for erically co ti g t e erivative of sol tio for a ilto -Jaco i eq atio s ( J s). e it is a lie for t e cro ot s i g ro le , it irectly e larges t e o for sta ilizatio so t at t e o ar ositio is i cl e i o . or a s rvey a ot er sol tio et o s for J s, e refer to g ilar a re er (2014); liy (2011); eeler et al. (2000); kes (1969); avasca a re er (2007).\ne orga izatio of t e a er is as follo s. e cro ot o el is i tro ce i \u00a7 2. \u00a7 3 s arizes t e t eory of t e sta le a ifol et o for J s. o troller esig is recisely ex lai e i \u00a7 5 a si latio res lts are s o i \u00a7 6.\n2. I I\nI t is sectio , e erive a o li ear o el of t e cro ot. igs. 1, 2 s o t e cro ot a its sc e atic o el. e co trol torq e is a lie o ly to t e seco joi t fro a act ator t ro g a lley a a ti i g elt.\nor t e it li k (i 1, 2), qi is t e a gle, i is t e ass, li is t e le gt , lci is t e ista ce fro it joi t to t e ce ter of ass ( ), i is t e i ertia aro t e ce ter of ass, a let g e a gravitatio al acceleratio (9.801\n10th IF y posiu on onlinear ontrol yste s ugust 23-25, 2016. onterey, alifornia,\nopyright 2016 IF 380\nSwing up and stabilization of the cr bot via nonlinear opti al control based on\nstable anifold ethod\nTakamasa Horibe \u2217 Noboru Sakamoto \u2217\u2217\n\u2217 Department of Aerospace Engineering, Grad ate School f Engineering, Nagoya University, Furo-cho, Chikus ku, Nagoya,\nJapan, (e-mail: oribe.takamasa@j.mbox.nagoya-u.ac.jp) \u2217\u2217 Dep rtment of Mech tronics, Faculty of Science and Engineeri g, Nanzan University, Yamazato-ch 18, Showa ku, Nagoya, Japan,\n(e-mail: noboru.sakamoto@nanzan-u.ac.jp)\nAbstract: This paper c nsiders the problem of swing up and stabilization for the Acrobot. It is shown that st manifold metho which has b en proposed f r computing nonline r optimal is capable of designing f controllers for this problem. An optimal stabiliz con roller is obt i ed as a single feedback law by n merically solving a Hamilton-Jacobi equation by the stable manifold method. It s shown that u like existing methods for Acrobot swi g up uch as par ial feedback linearization, the resultant control is me hanically digenous in the sense that it uses reactions of arms ly and, as a cons quence, control input is kept low. A number of simulations verify the effectiveness and robustness of the controller.\nK yw rds: Acrobot, Nonlinear optimal control, Hamilton-Jacobi equation, stable manifold method\n1. INTRODUCTION\nUnderactuated systems are mechanical systems which have ewer con rol inputs than degrees of freedom. Control of underactuat d systems is current y n active topic for many esearchers due to wide application range in Robotics or aerospace field (Liu and Yu, 2013; Xin an Liu, 2013). The Acrobot is a 2-dimensional underactuated mechanical system of en us d as a benchmark problem for testing nonlinea control methods. I consists of two links and an a tuat r is installed only at the ec nd join . A co mon control objective of he Acrobot is to swing it up fr m the downward position to the un t ble upri ht po ition and to stabilize it vertically. This is a challe ging task because of the movement in a large range of nonlinearity.\nGenerally, the swing up control is divided into tw phases first, the swing up phase in whic nonlinearity is dominant, and then, st bilizati n phase which estabishes autonomus s ability in a neighborhood of the origin. The e are some other effective ways to design swing up c trollers such as using p rtial feedback linearization (Spong, 1995), energy feedba k (Xin nd Yamas ki, 2012; Xin and Kaneda, 2007), trajectory t acking (Zhang et al., 2013), Lyapunov based control (Zergeroglu et al., 1998) and intellige t (Brown and Passino, 1997). However, switching contr ller has no guarantee of stability in the vicinity of the boundary. Researchers in (Davison and Bortoff, 1997; Xin and Kaneda, 2001) pr pose methods to enlarge the region of attraction (RoA) of linear controllers, s abiliza tion of the Acrobot by linear control is inherently ifficult. Backstepping approachOlfati-Saber (2000) should be\nmentioned as a singl feedback control method under some assumptions that are difficult to satisfy generally.\nIn this paper, we show that it is possible to design a single (without witching) optimal feedback controller for swing up and s abilization of the Acrobot using the stable manif ld method (Sakamoto and van der Schaft, 2008; Sakamoto, 2013). The method has been developed for numerically computi g the derivative of solution Hamilton-Jacobi equations (HJEs). Wh n it is applied for the Acrobot swing up problem, it directly enlarges the RoA for stabilization so that the downward position i included in RoA. For a survey and other solution methods for HJEs, w r fer to Aguilar and Krener (2014); Aliyu (2011); Beeler et al. (2000); Lukes (1969); Navasca and Krener (2007).\nThe organization of the paper i as follows. The Acrob t model is introduced in \u00a7 2. \u00a7 3 summarizes th th ory of th stable manifold method for HJEs. Controller de ign is precisely explained in \u00a7 5 and simulation results are shown in \u00a7 6.\n2. MODELING AND ANALYSIS OF THE ACROBOT\nIn this section, we derive a nonlinear model of the Acrobot. Figs. 1, 2 show the Acrob t and i s schematic model. The control torque is ap lied only to he second joint from an actuator through a pulley and a timing belt.\nFor the ith link (i = 1, 2), qi is the angle, mi is the mass, li is the length, lci is the distance from ith joint to the center of mass (COM), Ji is the inertia around the center of mass, and let g be a gravitational acceleration (9.801\n10th IFAC Symposium on Nonlinear Control Systems August 23-25, 2016. Monterey, California, USA\nCopyright \u00a9 2016 IFAC 380",
            "Takamasa Horibe et al. / IFAC-PapersOnLine 49-18 (2016) 374\u2013379 375\nSwing up and stabilization of the Acrobot via nonlinear optimal control based on\nstable manifold method\nTakamasa Horibe \u2217 Noboru Sakamoto \u2217\u2217\n\u2217 Department of Aerospace Engineering, Graduate School of Engineering, Nagoya University, Furo-cho, Chikusa-ku, Nagoya,\nJapan, (e-mail: horibe.takamasa@j.mbox.nagoya-u.ac.jp) \u2217\u2217 Department of Mechatronics, Faculty of Science and Engineering, Nanzan University, Yamazato-cho 18, Showa-ku, Nagoya, Japan,\n(e-mail: noboru.sakamoto@nanzan-u.ac.jp)\nAbstract: This paper considers the problem of swing up and stabilization for the Acrobot. It is shown that stable manifold method which has been proposed for computing nonlinear optimal control is capable of designing feedback controllers for this problem. An optimal stabilization controller is obtained as a single feedback law by numerically solving a Hamilton-Jacobi equation by the stable manifold method. It is shown that unlike existing methods for Acrobot swing up such as partial feedback linearization, the resultant control is mechanically indigenous in the sense that it uses reactions of arms effectively and, as a consequence, control input is kept low. A number of simulations verify the effectiveness and robustness of the controller.\nKeywords: Acrobot, Nonlinear optimal control, Hamilton-Jacobi equation, stable manifold method\n1. INTRODUCTION\nUnderactuated systems are mechanical systems which have fewer control inputs than degrees of freedom. Control of underactuated systems is currently an active topic for many researchers due to wide application range in Robotics or aerospace field (Liu and Yu, 2013; Xin and Liu, 2013). The Acrobot is a 2-dimensional underactuated mechanical system often used as a benchmark problem for testing nonlinear control methods. It consists of two links and an actuator is installed only at the second joint. A common control objective of the Acrobot is to swing it up from the downward position to the unstable upright position and to stabilize it vertically. This is a challenging task because of the movement in a large range of nonlinearity.\nGenerally, the swing up control is divided into two phases, first, the swing up phase in which nonlinearity is dominant, and then, stabilization phase which estabishes autonomus stability in a neighborhood of the origin. There are some other effective ways to design swing up controllers such as using partial feedback linearization (Spong, 1995), energy feedback (Xin and Yamasaki, 2012; Xin and Kaneda, 2007), trajectory tracking (Zhang et al., 2013), Lyapunov based control (Zergeroglu et al., 1998) and intelligent control (Brown and Passino, 1997). However, switching controller has no guarantee of stability in the vicinity of the boundary. Researchers in (Davison and Bortoff, 1997; Xin and Kaneda, 2001) propose methods to enlarge the region of attraction (RoA) of linear controllers, stabilization of the Acrobot by linear control is inherently difficult. Backstepping approachOlfati-Saber (2000) should be\nmentioned as a single feedback control method under some assumptions that are difficult to satisfy generally.\nIn this paper, we show that it is possible to design a single (without switching) optimal feedback controller for swing up and stabilization of the Acrobot using the stable manifold method (Sakamoto and van der Schaft, 2008; Sakamoto, 2013). The method has been developed for numerically computing the derivative of solution for Hamilton-Jacobi equations (HJEs). When it is applied for the Acrobot swing up problem, it directly enlarges the RoA for stabilization so that the downward position is included in RoA. For a survey and other solution methods for HJEs, we refer to Aguilar and Krener (2014); Aliyu (2011); Beeler et al. (2000); Lukes (1969); Navasca and Krener (2007).\nThe organization of the paper is as follows. The Acrobot model is introduced in \u00a7 2. \u00a7 3 summarizes the theory of the stable manifold method for HJEs. Controller design is precisely explained in \u00a7 5 and simulation results are shown in \u00a7 6.\n2. MODELING AND ANALYSIS OF THE ACROBOT\nIn this section, we derive a nonlinear model of the Acrobot. Figs. 1, 2 show the Acrobot and its schematic model. The control torque is applied only to the second joint from an actuator through a pulley and a timing belt.\nFor the ith link (i = 1, 2), qi is the angle, mi is the mass, li is the length, lci is the distance from ith joint to the center of mass (COM), Ji is the inertia around the center of mass, and let g be a gravitational acceleration (9.801\n10th IFAC Symposium on Nonlinear Control Systems August 23-25, 2016. Monterey, California, USA\nCopyright \u00a9 2016 IFAC 380\nSwing up and stabilization of the Acrobot via nonlinear optimal control based on\nstable manifold method\nTakamasa Horibe \u2217 Noboru Sakamoto \u2217\u2217\n\u2217 Department of Aerospace Engineering, Graduate School of Engineering, Nagoya University, Furo-cho, Chik sa-ku, Nag ya,\nJapan, (e-mail: horibe.takamasa@j.mbox.nagoy u.ac.jp) \u2217\u2217 Department of Mec atronics, Faculty of Science and Engineering, Nanz n University, Y mazato-cho 18, Showa-ku, Nagoya, Japa ,\n(e-mail: noboru.sakam to@nanzan u.ac.jp)\nAbstract: This paper considers the problem of swing up and stabilization for the Acrobot. It is shown that stable manif ld method which has been proposed for computing nonlinear optimal control is cap of designing fee back controll rs for this pr blem. An optimal st bilization ler is obtained as a single f law by numerically solving a Hamilton-Jacobi equ by he stable m ifold method. It is shown that nlike existing methods for Acrobot swing up such as partial feedback linearizat on, the resulta t control is mechanically indigenous i the ense that i uses reactions of arms effectively and, as a consequen e, control put is kept low. A number of simulations verify the ness and robustn ss of the controller.\nKeywords: Acrobot, Nonlinear optimal control, Hamilton-Jacobi equation, stable manifold m th d\n1. INTRODUCTION\nUnderactuated systems are mechanical systems which have fewer control inputs than degrees of freedom. Control o underac uated systems is currently an active topic for many research rs due to wide app ic tion range in Robotics o aerospace field (Liu and Yu, 2013; Xin and Liu, 2013). The Acrobot is a 2-dimensional underactuate mechanical system often used as a benchmark problem for testing nonlinear con rol m thods. It consists of two links and an actuato is installed only a the second joint. A common ontr l objective of the Acrobot i t swing i up fro the downward position o the unstable upright positi n and to stabilize it vertically. This i challen ing ta k because of the movement in a large range of no linearity.\nGenerally, the swing up control is divided into two phases, first, the swing up phase in which nonlinearity is d minant and then, stabilization phase w ich estabishes autonomus stability in neighb rhood of the origin. There are some o her effective ways to design swing up cont ollers such as using partial feedback linearization (Sp g, 1995), energy feedb ck (Xin and Yamasaki, 2012; Xin and Kaneda, 2007), traje tory tr cking (Zh ng et al., 2013), Lyapunov based control (Ze geroglu et al., 1998) and intelligent control (Brown and Passino, 1997). However, switchi g ler has no guarantee of stability in the vicinity of the b undary. Researchers in (Davison and Bortoff, 1997; Xin and Kaneda, 2001) propose methods to enlarge the region of attraction (RoA) f linear controllers, stabilization of the Acrobot by linear control is inheren ly diffi cult. Backstepping approachOlfati-Saber (2000) shoul be\nmentioned as a single feedback control method under some assumptions that ar difficult to satisfy generally.\nIn this paper, we show that it is possible to design a single (without switching) optimal feedback controller for swing up and tabilization of the Acrobot using the stable manifold me hod (Sakamoto and van der Schaft, 2008; Sakam to, 2013). The method has been developed for numerically computing the derivative of solution for Hamilton-Jacobi equatio s (HJEs). When it is applied the Acrobot swing up problem, it dir ctly enlarges the RoA for stabilization so that the downward position is included in RoA. For a survey and other solution method for HJEs, we refer to Aguilar and Krener (2014); Aliyu (2011); Beel r t al. (2000); Lukes (1969); Navasca and Krener (2007).\nThe organization of the paper is as follows. The Acrobot model is introduced in \u00a7 2. \u00a7 3 ummarizes the theory f the stable manifold method for HJEs. Controll r d sign is pr cisely explained in \u00a7 5 and simulation results are hown in \u00a7 6.\n2. MODELING AND ANALYSIS OF THE ACROBOT\nIn this section, we derive a nonlinear model of the Acrobot. Figs. 1, 2 show the Acrobot and its schematic model. The control torque is applied nly to he second joint from an actuator through a ulley and a iming belt.\nFor the ith link (i = 1, 2), qi is the angle, mi is the mass, li is the length, lci is the distance from ith joint to the center of mass (COM), Ji is the inertia around the center of mass, and let g be a gravitational acceleration (9.801\n10th IFAC Symposium on Nonlinear Control Systems August 23-25, 2016. Monterey, California, USA\nCopyright \u00a9 2016 IFAC 380\nS ing up and stabilization of the crobot via nonlinear opti al control based on\nstable anifold ethod\nTakamasa Horibe \u2217 Noboru Sakamoto \u2217\u2217\n\u2217 Department of Aerospace Engineering, Graduate School of Engineering, Nagoya University, Furo-cho, Chikusa-ku, Nagoya,\nJapan, (e-mail: horibe.takamasa@j.mbox.nagoya-u.ac.jp) \u2217\u2217 Department of Mechatronics, Faculty of Science and Engineering, Nanzan University, Yamazato-cho 18, Showa-ku, Nagoya, Japan,\n(e-mail: noboru.sakamoto@nanzan-u.ac.jp)\nAbstract: This paper considers the problem of swing up and stabilization for the Acrobot. It is shown that stable manifold method which has been proposed for computing nonlinear optimal control is capable of designing feedback controllers for this problem. An optimal stabilization controller is obtained as a single feedback law by numerically solving a Hamilton-Jacobi equation by the stable manifold method. It is shown that unlike existing methods for Acrobot swing up such as partial feedback linearization, the resultant control is mechanically indigenous in the sense that it uses reactions of arms effectively and, as a consequence, control input is kept low. A number of simulations verify the effectiveness and robustness of the controller.\nKeywords: Acrobot, Nonlinear optimal control, Hamilton-Jacobi equation, stable manifold method\n1. INTRODUCTION\nUnderactuated systems are mechanical systems which have fewer control inputs than degrees of freedom. Control of underactuated systems is currently an active topic for many researchers due to wide application range in Robotics or aerospace field (Liu and Yu, 2013; Xin and Liu, 2013). The Acrobot is a 2-dimensional underactuated mechanical system often used as a benchmark problem for testing nonlinear control methods. It consists of two links and an actuator is installed only at the second joint. A common control objective of the Acrobot is to swing it up from the downward position to the unstable upright position and to stabilize it vertically. This is a challenging task because of the movement in a large range of nonlinearity.\nGenerally, the swing up control is divided into two phases, first, the swing up phase in which nonlinearity is dominant, and then, stabilization phase which estabishes autonomus stability in a neighborhood of the origin. There are some other effective ways to design swing up controllers such as using partial feedback linearization (Spong, 1995), energy feedback (Xin and Yamasaki, 2012; Xin and Kaneda, 2007), trajectory tracking (Zhang et al., 2013), Lyapunov based control (Zergeroglu et al., 1998) and intelligent control (Brown and Passino, 1997). However, switching controller has no guarantee of stability in the vicinity of the boundary. Researchers in (Davison and Bortoff, 1997; Xin and Kaneda, 2001) propose methods to enlarge the region of attraction (RoA) of linear controllers, stabilization of the Acrobot by linear control is inherently difficult. Backstepping approachOlfati-Saber (2000) should be\nmentioned as a single feedback control method under some assumptions that are difficult to satisfy generally.\nIn this paper, we show that it is possible to design a single (without switching) optimal feedback controller for swing up and stabilization of the Acrobot using the stable manifold method (Sakamoto and van der Schaft, 2008; Sakamoto, 2013). The method has been developed for numerically computing the derivative of solution for Hamilton-Jacobi equations (HJEs). When it is applied for the Acrobot swing up problem, it directly enlarges the RoA for stabilization so that the downward position is included in RoA. For a survey and other solution methods for HJEs, we refer to Aguilar and Krener (2014); Aliyu (2011); Beeler et al. (2000); Lukes (1969); Navasca and Krener (2007).\nThe organization of the paper is as follows. The Acrobot model is introduced in \u00a7 2. \u00a7 3 summarizes the theory of the stable manifold method for HJEs. Controller design is precisely explained in \u00a7 5 and simulation results are shown in \u00a7 6.\n2. MODELING AND ANALYSIS OF THE ACROBOT\nIn this section, we derive a nonlinear model of the Acrobot. Figs. 1, 2 show the Acrobot and its schematic model. The control torque is applied only to the second joint from an actuator through a pulley and a timing belt.\nFor the ith link (i = 1, 2), qi is the angle, mi is the mass, li is the length, lci is the distance from ith joint to the center of mass (COM), Ji is the inertia around the center of mass, and let g be a gravitational acceleration (9.801\n10th IFAC Symposium on Nonlinear Control Systems August 23-25, 2016. Monterey, California, USA\nCopyright \u00a9 2016 IFAC 380\ni ili i f i li i l l\nl if l\ns ri e \u2217 r t \u2217\u2217\n\u2217 epart e t of erospace gi eeri g, rad ate chool of gi eeri g, agoya iversity, ro-cho, hik sa-k , agoya, Japa , (e- ail: horibe.taka asa j. box. agoya- .ac.jp) \u2217\u2217 epart e t of echatro ics, ac lty of cie ce a d gi eeri g, a za iversity, a azato-cho 18, ho a-k , agoya, Japa ,\n(e- ail: obor .saka oto a za - .ac.jp)\nstr ct: is a er co si ers t e ro le of s i g a sta ilizatio for t e cro ot. It is s o t at sta le a ifol et o ic as ee ro ose for co ti g o li ear o ti al co trol is ca a le of esig i g fee ack co trollers for t is ro le . o ti al sta ilizatio co troller is o tai e as a si gle fee ack la y erically solvi g a a ilto -Jaco i eq atio y t e sta le a ifol et o . It is s o t at like existi g et o s for cro ot s i g s c as artial fee ack li earizatio , t e res lta t co trol is ec a ically i ige o s i t e se se t at it ses reactio s of ar s e ectively a , as a co seq e ce, co trol i t is ke t lo .\ner of si latio s verify t e e ective ess a ro st ess of t e co troller.\ney ords: cro ot, o li ear o ti al co trol, a ilto -Jaco i eq atio , sta le a ifol et o\n1. I I\neract ate syste s are ec a ical syste s ic ave fe er co trol i ts t a egrees of free o . o - trol of eract ate syste s is c rre tly a active to ic for a y researc ers e to i e a licatio ra ge i o otics or aeros ace el ( i a , 2013; i a i , 2013). e cro ot is a 2- i e sio al eract ate ec a ical syste ofte se as a e c ark ro le for testi g o li ear co trol et o s. It co sists of t o li ks a a act ator is i stalle o ly at t e seco joi t. co o co trol o jective of t e cro ot is to s i g it fro t e o ar ositio to t e sta le rig t ositio a to sta ilize it vertically. is is a c alle gi g task eca se of t e ove e t i a large ra ge of o li earity.\ne erally, t e s i g co trol is ivi e i to t o ases, rst, t e s i g ase i ic o li earity is o i a t, a t e , sta ilizatio ase ic esta is es a to o s sta ility i a eig or oo of t e origi . ere are so e ot er e ective ays to esig s i g co trollers s c as si g artial fee ack li earizatio ( o g, 1995), e - ergy fee ack ( i a a asaki, 2012; i a a e a, 2007), trajectory tracki g ( a g et al., 2013), ya ov ase co trol ( ergerogl et al., 1998) a i tellige t co trol ( ro a assi o, 1997). o ever, s itc i g co troller as o g ara tee of sta ility i t e vici ity of t e o ary. esearc ers i ( aviso a orto , 1997; i a a e a, 2001) ro ose et o s to e large t e regio of attractio ( o ) of li ear co trollers, sta ilizatio of t e cro ot y li ear co trol is i ere tly i - c lt. ackste i g a roac lfati- a er (2000) s o l e\ne tio e as a si gle fee ack co trol et o er so e ass tio s t at are i c lt to satisfy ge erally.\nI t is a er, e s o t at it is ossi le to esig a si gle ( it o t s itc i g) o ti al fee ack co troller for s i g a sta ilizatio of t e cro ot si g t e sta le a ifol et o ( aka oto a va er c aft, 2008; aka oto, 2013). e et o as ee evelo e for erically co ti g t e erivative of sol tio for a ilto -Jaco i eq atio s ( J s). e it is a lie for t e cro ot s i g ro le , it irectly e larges t e o for sta ilizatio so t at t e o ar ositio is i cl e i o . or a s rvey a ot er sol tio et o s for J s, e refer to g ilar a re er (2014); liy (2011); eeler et al. (2000); kes (1969); avasca a re er (2007).\ne orga izatio of t e a er is as follo s. e cro ot o el is i tro ce i \u00a7 2. \u00a7 3 s arizes t e t eory of t e sta le a ifol et o for J s. o troller esig is recisely ex lai e i \u00a7 5 a si latio res lts are s o i \u00a7 6.\n2. I I\nI t is sectio , e erive a o li ear o el of t e cro ot. igs. 1, 2 s o t e cro ot a its sc e atic o el. e co trol torq e is a lie o ly to t e seco joi t fro a act ator t ro g a lley a a ti i g elt.\nor t e it li k (i 1, 2), qi is t e a gle, i is t e ass, li is t e le gt , lci is t e ista ce fro it joi t to t e ce ter of ass ( ), i is t e i ertia aro t e ce ter of ass, a let g e a gravitatio al acceleratio (9.801\n10th IF y posiu on onlinear ontrol yste s ugust 23-25, 2016. onterey, alifornia,\nopyright 2016 IF 380\nSwing up and stabilization of the cr bot via nonlinear opti al control based on\nstable anifold ethod\nTakamasa Horibe \u2217 Noboru Sakamoto \u2217\u2217\n\u2217 Department of Aerospace Engineering, Grad ate School f Engineering, Nagoya University, Furo-cho, Chikus ku, Nagoya,\nJapan, (e-mail: oribe.takamasa@j.mbox.nagoya-u.ac.jp) \u2217\u2217 Dep rtment of Mech tronics, Faculty of Science and Engineeri g, Nanzan University, Yamazato-ch 18, Showa ku, Nagoya, Japan,\n(e-mail: noboru.sakamoto@nanzan-u.ac.jp)\nAbstract: This paper c nsiders the problem of swing up and stabilization for the Acrobot. It is shown that st manifold metho which has b en proposed f r computing nonline r optimal is capable of designing f controllers for this problem. An optimal stabiliz con roller is obt i ed as a single feedback law by n merically solving a Hamilton-Jacobi equation by the stable manifold method. It s shown that u like existing methods for Acrobot swi g up uch as par ial feedback linearization, the resultant control is me hanically digenous in the sense that it uses reactions of arms ly and, as a cons quence, control input is kept low. A number of simulations verify the effectiveness and robustness of the controller.\nK yw rds: Acrobot, Nonlinear optimal control, Hamilton-Jacobi equation, stable manifold method\n1. INTRODUCTION\nUnderactuated systems are mechanical systems which have ewer con rol inputs than degrees of freedom. Control of underactuat d systems is current y n active topic for many esearchers due to wide application range in Robotics or aerospace field (Liu and Yu, 2013; Xin an Liu, 2013). The Acrobot is a 2-dimensional underactuated mechanical system of en us d as a benchmark problem for testing nonlinea control methods. I consists of two links and an a tuat r is installed only at the ec nd join . A co mon control objective of he Acrobot is to swing it up fr m the downward position to the un t ble upri ht po ition and to stabilize it vertically. This is a challe ging task because of the movement in a large range of nonlinearity.\nGenerally, the swing up control is divided into tw phases first, the swing up phase in whic nonlinearity is dominant, and then, st bilizati n phase which estabishes autonomus s ability in a neighborhood of the origin. The e are some other effective ways to design swing up c trollers such as using p rtial feedback linearization (Spong, 1995), energy feedba k (Xin nd Yamas ki, 2012; Xin and Kaneda, 2007), trajectory t acking (Zhang et al., 2013), Lyapunov based control (Zergeroglu et al., 1998) and intellige t (Brown and Passino, 1997). However, switching contr ller has no guarantee of stability in the vicinity of the boundary. Researchers in (Davison and Bortoff, 1997; Xin and Kaneda, 2001) pr pose methods to enlarge the region of attraction (RoA) of linear controllers, s abiliza tion of the Acrobot by linear control is inherently ifficult. Backstepping approachOlfati-Saber (2000) should be\nmentioned as a singl feedback control method under some assumptions that are difficult to satisfy generally.\nIn this paper, we show that it is possible to design a single (without witching) optimal feedback controller for swing up and s abilization of the Acrobot using the stable manif ld method (Sakamoto and van der Schaft, 2008; Sakamoto, 2013). The method has been developed for numerically computi g the derivative of solution Hamilton-Jacobi equations (HJEs). Wh n it is applied for the Acrobot swing up problem, it directly enlarges the RoA for stabilization so that the downward position i included in RoA. For a survey and other solution methods for HJEs, w r fer to Aguilar and Krener (2014); Aliyu (2011); Beeler et al. (2000); Lukes (1969); Navasca and Krener (2007).\nThe organization of the paper i as follows. The Acrob t model is introduced in \u00a7 2. \u00a7 3 summarizes th th ory of th stable manifold method for HJEs. Controller de ign is precisely explained in \u00a7 5 and simulation results are shown in \u00a7 6.\n2. MODELING AND ANALYSIS OF THE ACROBOT\nIn this section, we derive a nonlinear model of the Acrobot. Figs. 1, 2 show the Acrob t and i s schematic model. The control torque is ap lied only to he second joint from an actuator through a pulley and a timing belt.\nFor the ith link (i = 1, 2), qi is the angle, mi is the mass, li is the length, lci is the distance from ith joint to the center of mass (COM), Ji is the inertia around the center of mass, and let g be a gravitational acceleration (9.801\n10th IFAC Symposium on Nonlinear Control Systems August 23-25, 2016. Monterey, California, USA\nCopyright \u00a9 2016 IFAC 380\nm/s2). The equation of motion of the Acrobot is derived as follows by the method of Lagrange.\nM(q2)q\u0308 + C(q, q\u0307)q\u0307 +G(q) = \u03c4, (1)\nwhere\nM(q2) = [ M11(q2) M12(q2) M21(q2) M22 ]\n=\n[ a1 + a2 + 2a3 cos q2 a2 + a3 cos q2\na2 + a3 cos q2 a2\n]\nC(q, q\u0307) = [ \u2212a3q\u03072 sin q2 \u2212a3(q\u03071 + q\u03072) sin q2 a3q\u03071 sin q2 0 ] ,\nG(q) =\n[ \u2212b1 sin q1 \u2212 b2 sin(q1 + q2)\n\u2212b2 sin(q1 + q2)\n] ,\na1 = m1L 2 c1 +m2L 2 1 + J1, a2 = m2L 2 c2 + J2, a3 = m2L1Lc2, b1 = (m1Lc1 +m2L1)g, b2 = m2Lc2g.\n\u03c4 = [\u03c41, \u03c42] T are the torque on the joints. We assume that there exists resistance force proportional with angular velocity such as viscous friction or counter electromotive force of the actuator. Then \u03c4 is given as\n\u03c41 = \u2212\u00b51q\u03071, \u03c42 = nKDCu\u2212 \u00b52q\u03072,\nwhere u is the control input voltage for the actuator, n is gear ratio of the pulley, KDC is electromotive torque constant and \u00b51, \u00b52 is the viscous resistance coefficient of the joints.\nDefining x = [x1, x2, x3, x4] T = [q1, q2, q\u03071, q\u03072] T as system variables, and letting H(q, q\u0307) = [H1(q, q\u0307), H2(q, q\u0307)]\nT = C(q, q\u0307)q\u0307+G(q)+ [\u00b51q\u03071, \u00b52q\u03072]\nT , the dynamic equation (1) is rewritten in the state space form as\nx\u0307 = f(x) + g(x)u, (2)\nwhere\nf(x) =  \nx3 x4\n\u2212M\u22121(q2) [ H1(q, q\u0307) H2(q, q\u0307)\n]   ,\ng(x) =  \n0 0\n\u2212M\u22121(q2)\n[ 0\nnKDC\n]   .\nThe purpose of this paper is to design a single nonlinear optimal feedback controller for swinging up and stabiliz-\ning the Acrobot from the stable equilibrium point x = [\u03c0, 0, 0, 0] to the unstable equilibrium point x = [0, 0, 0, 0]. The system parameters are shown in Table.1.\n3. HAMILTON-JACOBI EQUATION AND STABLE MANIFOLD\nIn this section, we briefly review the stable manifold method proposed in Sakamoto and van der Schaft (2008); Sakamoto (2013) for numerically solving HJEs for nonlinear optimal control problem. It is first shown that a stable manifold of the Hamiltonian system associated with a HJE is equivalent to the stabilizing solution of the HJE (I). Then, the stable manifold method algorithm to compute flows on the stable manifold of the Hamiltonian system is presented (II). Finally, the optimal state feedback function is constructed from the flow data by polynomial functions that define the stable manifold (III). In what follows, the procedures (I)\u223c(III) will be described with some details.\nLet us consider the optimal regulation problem for the following nonlinear affine system and quadratic cost function J . \n\nx\u0307 = f(x) + g(x)u, x(0) = x0\nJ = 1\n2\n\u222b \u221e\n0\n( xTQx+ uTRu ) dt,\n(3)\nwhere x \u2208 Rn, u \u2208 U \u2208 Rm and R \u2208 Rn\u00d7m, Q \u2208 Rn\u00d7m are positive-definite matrices. We also assume that f(\u00b7) : Rn \u2192 Rn with f(0) = 0 and g(\u00b7) : Rn \u2192 R are all C\u221e. It is then possible to write f(x) = Ax+o(|x|), g(x) = B+O(|x|) with real matrices A \u2208 Rn\u00d7n, B \u2208 Rn\u00d7m.\nIFAC NOLCOS 2016 August 23-25, 2016. Monterey, California, USA",
            "376 Takamasa Horibe et al. / IFAC-PapersOnLine 49-18 (2016) 374\u2013379\nDenote R\u0304(x) = g(x)R\u22121g(x)T . Applying dynamic programming to the above problem, an HJE is derived as\n(HJ) H(x, p) = pT f(x)\u2212 1\n2 pT R\u0304(x)p+\n1 2 xTQx = 0\nwhere p1 = \u2202V/\u2202x1, \u00b7 \u00b7 \u00b7 , pn = \u2202V/\u2202xn.\nThe stabilization solution of (HJ) is defined as follows.\nDefinition 1. A solution V (x) of (HJ) is said to be the stabilization solution if x = 0 is an asymptotically stable equilibrium of a vector field \u2202H\n\u2202p (x, p(x)), where p(x) =\n(\u2202V/\u2202x)T (x).\n(I) A sufficient condition for the local stabilizing solution for (HJ) is obtained in van der Schaft (1991). It is a natural condition based on a linearization argument. We recall that, for the algebraic Riccati equation\n(RIC) PA+ATP \u2212 P \u00afR(0)P +Q = 0,\nwhich is the linearization of (HJ), a symmetric matrix P is said to be the stabilizing solution of (RIC) if it is a solution of (RIC) and A\u2212R(0)P is stable.\nTheorem 1. If the Riccati equation has the stabilizing solution P , there exists, locally around the origin, the stabilizing solution V (x) to (HJ) with (\u22022V/\u2202x2)(0) = P . Moreover, the set {(x, p) | p = (\u2202V/\u2202x)T } is a stable manifold of\nx\u0307 = \u2202H \u2202p (x, p), p\u0307 = \u2212\u2202H \u2202x (x, p). (4)\nThroughout the paper, we assume that the stabilizing solution P to (RIC), which is a linearization of (HJ), exists. It means, as is stated in Theorem 1, that an n-dimensional stable manifold {(x, p)|p = p(x)} exists for (4) around the origin. Note that the function p(x) is used in feedback control. By a suitable linear coordinate transformation with T \u2208 Rn\u00d7n, the linear part of Hamiltonian system (4) is diagonalized to get[ x\u0307\u2032\np\u0307\u2032\n] = [ A\u2212 R\u0304P 0\n0 \u2212(A\u2212 R\u0304P )T\n] [ x\u2032\np\u2032\n] +higher order terms.\n(5) Recalling that P is the stabilizing solution of (RIC), A \u2212 R\u0304P is an asymptotically stable n\u00d7 n matrix.\n(II) The stable manifold algorithm is to compute flows on a stable manifold of a system of ordinary differential equations of the form\n\u03a3 :\n{ x\u0307 = Fx+ nx(x, p)\np\u0307 = \u2212FT p+ np(x, p) (6)\nwhere nx, np consist of high order nonlinear C1 functions and F \u2208 Rn\u00d7n is a stable matrix. Let us define the sequence {xk(t, \u03be)} and {pk(t, \u03be)} by\n\nxk+1 = eFt\u03be +\n\u222b t\n0\neF (t\u2212s)nx(xk(s), pk(s)) ds\npk+1 = \u2212 \u222b \u221e\nt\ne\u2212FT (t\u2212s)np(xk(s), pk(s)) ds\n(7)\nx0 = eFt\u03be, p0 = 0 (8)\nwith k = 0, 1, 2, \u00b7 \u00b7 \u00b7 and arbitrary \u03be \u2208 Rn. Then, the following theorem holds.\nTheorem 2. (Sakamoto and van der Schaft, 2008) The sequence xk(t, \u03be) and pk(t, \u03be) are convergent to zero for sufficiently small |\u03be|, that is, xk(t, \u03be), pk(t, \u03be) \u2192 0 as\nt \u2192 \u221e for all k = 0, 1, 2, \u00b7 \u00b7 \u00b7 . Furthermore, xk(t, \u03be) and pk(t, \u03be) are uniformly convergent to a solution of (6) on [0,\u221e) as k \u2192 \u221e. Let x(t, \u03be) and p(t, \u03be) be the limits of xk(t, \u03be) and pk(t, \u03be), respectively. Then, x(t, \u03be) and p(t, \u03be) are the solution on the stable manifold of (6), that is, x(t, \u03be), p(t, \u03be) \u2192 0 as t \u2192 \u221e.\nThis theorem states that we can approximate flows on the stable manifold of (6) by the sequence (7).\n(III) We apply the algorithm in (7)-(8) to (5) to get x\u2032 k(t, \u03be), p \u2032 k(t, \u03be). The flows on the stable manifold in the original coordinates is given by[ xk(t, \u03be) pk(t, \u03be) ] = T [ x\u2032 k(t, \u03be) p\u2032k(t, \u03be) ]\nand the set\n\u2126k = {(xk(t, \u03be), pk(t, \u03be)) | |\u03be|:sufficiently small, t \u2208 R} is a parametrization of the stable manifold. The approximation accuracy can be easily improved by employing larger k. If one can rewrite \u2126k with the relation p = \u03c0k(x), then \u03c0k(x) will serve as an approximation of (\u2202V/\u2202x)(x). In our application for the Acrobot control, we will approximately obtain (\u2202V/\u2202x)(x) via functional fitting using polynomials.\n4. CONTROLLER DESIGN\nWe apply the stable manifold method to the Acrobot. The weight matrix Q and R in (3) are chosen as follows considering maximum input voltage and stability by the corresponding LQR in the linear domain.\nQ = diag (0.001, 0.1, 0.0005, 0.1) , R = 1\nThe closed-loop poles \u03bb of the system linearized around the origin with this weighting matrices are\n\u03bb = \u221213.37, \u22125.35, \u22124.92, \u22124.64.\nIn this approach, the trajectory xk(t, \u03be) on the stable manifold is numerically computed for a large number of \u03be\u2019s. We have successfully found a \u03be from which the corresponding function xk(t, \u03be) passes through the pending position (\u03c0, 0, 0, 0). We have also found a different \u03be for which trajectory xk(t, \u03be) reaches the pending position with a much longer and involved trajectory. These two closed loop trajectories are show in Fig. 3, blue line (trajectory 1) and red dotted line (trajectory 2). They are on the same stable manifold and they both satisfy the same HJE for the optimal control problem. This phenomena is know to be an issue of non-uniqueness of solution of a HJE (Day, 1998) and has been firstly observed in a physical system in Fujimoto and Sakamoto (2011).\nNote that the stable manifold as an 4-dimensional manifold is unique but when it is projected on the base space x, a multi-valued function can be obtained. To see which one of the controllers is optimal, one has to compute the values of cost function. The values of the cost function are found to be 18.50(trajectory 1) and 28.01(trajectory 2) and one can see that the blue trajectory is optimal. The swing up motions by the two controllers are shown in Figs. 4, 5. In \u00a7 5, detailed simulations will be carried out for the optimal controller (trajectory 1).\nIFAC NOLCOS 2016 August 23-25, 2016. Monterey, California, USA"
        ]
    },
    {
        "image_filename": "designv11_14_0000158_j.mechmachtheory.2019.103625-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000158_j.mechmachtheory.2019.103625-Figure3-1.png",
        "caption": "Fig. 3. Two four-bar linkages within the linkage for IO analysis.",
        "texts": [
            " An input\u2013output (IO) equation is needed for position and branch analysis. The IO equation can be obtained using the forward kinematics of the 3-RRR parallel mechanism, which has been well documented in literature [26,27] . The reported methods are mainly based on the loop closure equation involving the input rotations and also the end-effector coordinate and orientation. In this work, we resort to the IO equation of four-bar linkages, in light of the nature of the 1-dof linkage, to take advan- tage of a well developed formulation [28] . Refer to Fig. 3 , there are two closed kinematic chains, namely, Loop 1 of B 1 C 1 A 1 A 2 C 2 B 2 and Loop 2 of B 2 C 2 A 2 A 3 C 3 B 3 . Let the input angle be \u03b8 . Here we drop the subscript of pose for clarity. The input angle at each limb is then equal to \u03b80 i + \u03b8, i = 1 , 2 , 3 . Thus the positions of C i are known once the input angle is given. We can then virtually construct two four-bar linkages out of the two loops, which are A 1 C 1 C 2 A 2 and A 2 C 2 C 3 A 3 . Two pairs of input\u2013output angles, { \u03c8 1 , \u03bc1 } and { \u03c8 2 , \u03bc2 }, are defined for them"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000876_0954407018824943-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000876_0954407018824943-Figure5-1.png",
        "caption": "Figure 5. Typical disposition of bearings supporting a shaft in a FEAD. FEAD: front engine accessory drive.",
        "texts": [
            "13 It can be fastidious but leads to a more accurate estimation of the total friction moment Mb. Alternatively, a simpler formulation can be used.19 Equation (10) represents the total bearing power losses in a FEAD PLbear = Xns i=1 PLb i = Xns i=1 PLb i F +PLb i R\u00f0 \u00de \u00f010\u00de where PLbear is the FEAD total bearing power loss; PLi b, the sum of power loss from the front PLb i F and rear PLb i R bearings of the ith shaft; and ns, the number of shafts composing the FEAD. It is worth noting that equation (10) suggests that there is at least a pair of bearings per shaft (Figure 5). Moreover, in the case of tensioner and idler-pulleys, the same theory is applied with L3=0 in Figure 5. The belt/pulley resultant radial force is one of the most important parameter of the SKF model.13 It depends on the system characteristics for which the bearing is operating. The radial force acting on each pulley is mainly the resultant force of the belt-span tensions around the pulley under analysis. To calculate this force for each pulley/shaft regardless of their positions in the FEAD, some vector operations on the considered forces are necessary. Indeed, by using the belt\u2013 pulley setting-unseating points s, u (calculated geometrically) along the FEAD path, one can obtain the distance vectors ~Di 1 and D ",
            " This may be done through dividing the distance vectors ~Di 1 and ~Di by their norms, equation (12) ~mTi 1 =~mDi 1 = ~Di 1 norm ~Di 1\u00f0 \u00de ~mTi =~mDi = ~Di norm ~Di\u00f0 \u00de 8< : \u00f012\u00de Next, the unit vectors ~mTi 1 and ~mTi shall be multi- plied by the scalars ~Ti 1 and ~Ti (equation (13)) which are the belt-span tensions obtained through an equilibrium analysis17 ~Ti 1 = j~Ti 1j~mTi 1 ~Ti = j~Tij~mTi ( \u00f013\u00de Finally, to provide the belt radial force Fri, which is the norm of ~Fri, needed for the SKF power loss model, a simple sum of vectors is performed, equation (14) ~Fr i =~Ti 1 +~Ti + ~Wi \u00f014\u00de Equation (14) is valid for bearings in the same plane of the pulleys (e.g. idler-pulleys) what, in practice, is not necessary is the case of all bearings in a FEAD. Very often, there is a shaft supported by a couple of bearings as in Figure 5. Consequently, the resultant radial force shall be calculated in each bearing. For example, let us consider the real case represented by the kth shaft in Figure 7. In this case, the forces used to calculate the bearing power losses are Fr k f and Fr k r which are deduced from Fr k using equilibrium equations of solid mechanics. Moreover, note that in Figure 7, since the pulley and the front bearing are in the same plane, L3 is nil. In contrast to belt internal losses (e.g. PLbelt hys) and in the class of belt external or mechanical losses, the speed loss Dv, then power loss DP is a characteristic of force conditioned (belt transmissions) systems"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001312_tie.2019.2952780-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001312_tie.2019.2952780-Figure8-1.png",
        "caption": "Fig. 8. Analysis of the stress and deformation of the Electromechanical actuator: (a) Deformation of the screw-threaded drawbar and (b) Safety factor for the screw-threaded drawbar.",
        "texts": [
            " The height of the spindle and the housing (F) is larger in the proposed model as it needs space for the planetary gear and spindle, which also makes the spindle height (G) bigger than the hydraulic model. Moreover, the dimensions are selected by considering the manufacturability of the electromechanical actuator as an independent module which can directly be installed on an existing lathe bed. During the simulation, the thrust of 50 kN is applied directly on the drawbar part, which is considered to be the weakest among the different parts of the electromechanical actuator. The deformation and the distribution of the safety factor are shown in the Fig. 8(a) and (b), respectively. From the FEA model, it is found that when 50 kN is applied, a deformation of 0.055 mm is produced in the drawbar. Moreover, based on the distribution of the safety factor and the considering the allowable stress of (material), a minimum safety factor of 6.62 can be selected for the drawbar. This safety factor is considered as the minimum safety factor for the whole actuator system during manufacturing process. Using the dimensions mentioned in Table II, a prototype of the electromechanical actuator is manufactured as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002922_j.mechmachtheory.2021.104382-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002922_j.mechmachtheory.2021.104382-Figure9-1.png",
        "caption": "Fig. 9. Pose of fully dynamic isotropy: (a) z = \u2212250 mm, (b) z = \u2212300 mm, and (c) z = \u2212350 mm.",
        "texts": [
            " Thereafter, the relationship between acceleration of the end-effector and driving torque can be written as \u03c4a = \u2212 G T a ( M p + 6 \u2211 i =1 J T Ci M i ) \u23a1 \u23a3 z\u0308 \u03b5 1 \u03b5 2 \u23a4 \u23a6 = B \u23a1 \u23a3 z\u0308 \u03b5 1 \u03b5 2 \u23a4 \u23a6 = B 1 \u0308z + B 23 [ \u03b5 1 \u03b5 2 ] (52) where B is a 3 \u00d7 3 matrix, B 1 is the first column of B , and B 23 is the second and third columns of B . Based on Eq. (36) , the dynamic isotropy analysis results of the parallel manipulator are shown in Fig. 8 . These shows that the performance of dynamic isotropy worsens near the singular locus. In addition, the dynamic isotropy performance improves when the z -coordinate decreases. The DDI is better when \u03d5 is 60 \u00b0, 180 \u00b0, or 300 \u00b0, which is consistent with the structural symmetry. In Fig. 8 (a)\u2013(c), when \u03b8 = 0 \u00b0, DDI = 1, and the corresponding poses are shown in Fig. 9 . These three poses are dynamically isotropic under the proposed evaluation method. It is clear that the parallel manipulator works in ideal poses, where the three limbs are symmetrical around the z -axis. After evaluating the performance of dynamic isotropy, the CAC of the 1T2R manipulator is analyzed. For the 1T2R parallel manipulator, the CAC is determined jointly by the three limbs. The specific analysis process is illustrated in Fig. 10 . First, the inequality group can be expressed geometrically, i",
            " When the z - coordinate becomes smaller (a longer stroke along the z -axis), the average CAC and the fluctuation of CAC in the orientation workspace improve simultaneously (i.e., GCAC becomes larger while FCAC gets smaller). With the help of the distributions of the DDI, CAC, GCAC, and FCAC, the suitable workspace can be identified according to the requirements of the application scenario. In addition, the working path can be optimized to achieve better dynamic performance. For the three poses shown in Fig. 9 , the dynamic conditioning indices (DCIs) [34] are 0.1578, 0.1578, and 0.1579, respectively. Under the evaluation of the DCI, DCI = 0 indicates that the manipulator is singular, and DCI = 1 means that the manipulator is dynamically isotropic. The three DCIs indicate that the three poses shown in Fig. 9 are away from dynamic isotropy. However, the three poses are ideal for this manipulator, and the dynamic performance of the three physical axes are exactly the same. Under the proposed evaluation method, the DDIs for the three poses are 1, which means that these poses are dynamically isotropic. In addition, the unit of the DCI for the 1T2R parallel manipulator is nonhomogeneous because of the mixed DoFs. Thus, the physical meaning of the DCI is ambiguous. Therefore, a dynamic performance evaluation from the aspect of driving torque is reasonable"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001019_rpj-07-2018-0182-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001019_rpj-07-2018-0182-Figure2-1.png",
        "caption": "Figure 2 Schematics of the experimental setup",
        "texts": [
            " Subsequently, another static analysis is done to analyze stress distribution before fabrication using FDM technique with ABS material. Hydroforming experimentation is done and the results are validated by comparing the convolution width formed during the experiment with FEA results. Metal bellows were hydroformed from tubes, which were produced by seamless welding of rolled sheets. Generally, for low quantity orders, convolutions are formed over these tubes successively using individual convolutionmethod (Witzenmann GmbH, 2010). Schematics of the experimental setup used for this study is shown in Figure 2. Existing die assembly consists of top and bottom die sets which are made up of right and left halves. Hydroforming a Metal bellow hydroforming Prithvirajan R. et al. Rapid Prototyping Journal D ow nl oa de d by N ot tin gh am T re nt U ni ve rs ity A t 0 2: 47 3 1 M ay 2 01 9 (P T ) single convolution involves four steps and they are repeated till the required number of convolution is formed. Step 1: The preformed metal tube is inserted over the rubber bladder and positioned at a required height where the convolution is to be formed (Figure 3)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002037_s00170-020-06366-8-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002037_s00170-020-06366-8-Figure12-1.png",
        "caption": "Fig. 12 a Printed inserts before removed from the build plate. b PBF-MSHAM printed mould inserts",
        "texts": [
            " In the printing with standard supports experiment, the lengths of time taken for preparing the print job, actual printing and support removal were 1 h, 56 h and 1 h, respectively. For simplicity reason, recorded time was rounded off to the nearest hour. In the printing with the MSHAM setup experiment, the lengths of time taken for preparing the print job and the actual printing were 3 h and 41 h, respectively. The support removal time was considered to be negligible as it only took a few minutes to remove the workpieces from the build plate. Figure 12a and b show the printed inserts before and after being removed from the build plate, respectively. The two sets of results were compiled and listed in Table 3. It can be seen that a reasonably big time saving of 14 h, or close to 25%, was recorded. The results showed that an extra 2 h was required in the MSHAM setup preparation because of the time taken to have the four base blocks pre-fabricated separately. The base substrate preparation time could vary depending on the complexity of the individual mould insert design"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001438_lra.2020.2969161-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001438_lra.2020.2969161-Figure4-1.png",
        "caption": "Fig. 4. Functional prototype: a slave device. (a) and (b) are the forceps-driver and the force sensor module embedded in the forceps driver. (c) is a schematic diagram of the force sensor. (d) indicates the force sensor calibration result. For the calibration, a commercial F/T sensor is utilized as a reference sensor.",
        "texts": [
            " The conceptually designed forceps-driver and the haptic feedback master device are fabricated as functional prototypes to carry out the experimental performance evaluation. In the experiment, the working characteristic and performances of each device, i.e., the forceps-driver and the master device, are separately investigated first. And then, the performance of the overall master-slave system is assessed to verify the applicability and usability of the proposed devices. The fabricated functional prototype of the slave device (i.e., the forceps-driver) is shown in Fig. 4(a). The overall dimension of the device is determined considering the size of a commercial microsurgical forceps (11200-14, Fine Science Tools). On the 3D-printed body, a micro servo motor (HV75K, MKS) and a potentiometer (3382H-12, BOURNS) are equipped to drive the rotator and to measure the angular displacement of the rotator, respectively. Fig. 4(b) indicate the fabricated tiny force sensor module applied to the forceps-driver. As schematically shown in Fig. 4(c), the soft PDMS is cast on the barometric sensor chip (MPL115A2, NXP) embedded PCB (printed circuit board) first, and then a low friction POM (polyoxymethylene) plate is bonded on the top of the cast PDMS block as a surface in contact with the forceps-handle. And, the fabricated sensor is calibrated with a commercial F/T (force/torque) sensor (Nano17, ATI). The calibration result shown in Fig. 4 (s) demonstrates that the force measurement capacity of the fabricated sensor is approximately 6.2 N. The two fabricated parts of the master device are shown in Fig. 5(a). The six diaphragm flexures are prototyped by lasercutting a POM plate. For the variable stiffness of the master device, the pre-tensioned wire-type SMA actuator of 0.1 mm in diameter (Flexinol, Dynalloy) is wound around the bearingpulleys that are arranged inside of the flexure mechanism. And, the magnet and the magnetic sensor PCB are assembled in the center inside each part to measure the displacement of push buttons that are produced by user\u2019s pinching motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000305_iros40897.2019.8968475-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000305_iros40897.2019.8968475-Figure2-1.png",
        "caption": "Fig. 2: Conventions for robot arm kinematic quantities as well as schematic of 2 real and 1 virtual accelerometer mounted on the arm",
        "texts": [
            " For experimental validations, a realistic dynamics model of a 7-DoF flexible joint manipulator is simulated and the results for different analysis are reported in Section IV. Similar experiments are performed for a 7-DoF flexible joint robot in Section V. Finally, the paper concludes in Section VI. In order to be able to estimate the joint acceleration, let us first consider an arbitrary link i of an articulated manipulator whose link side joint velocity q\u0307i \u2208 R and acceleration q\u0308i \u2208 R in joint i are unknown, see Fig. 2. The angular velocity and acceleration of link i are denoted by i\u03c9 \u2208 R3 and i\u03c9\u0307 \u2208 R3. Furthermore, a triaxial accelerometer mounted on link i measures Cartesian link acceleration ia \u2208 R3 in the accelerometer frame. The accelerometer frame is determined according to its sensing axis. Since the accelerometer location is fixed, the orientation of the sensor frame remains fixed with respect to the orientation of frame i, however, with a different origin. Due to noise and bias, the accelerometer signal needs to be fused with other sensors",
            " III-C). The offset ial is the linear acceleration caused by the joint frame translation. For computing ial (including gravity) of the i-th link frame, a virtual acceleration sensor (superscript v) located on joint i is introduced. Its output can be obtained with the help of data from the previous link as i\u22121a v = i\u22121aS1 + i\u22121\u03c9\u0307i\u22121 \u00d7 (i\u22121 Xv \u2212 i\u22121XS1 ) + i\u22121\u03c9i\u22121 \u00d7 ( i\u22121\u03c9i\u22121 \u00d7 (i\u22121 Xv \u2212 i\u22121XS1 )) , (4) where, i\u22121Xv is the position of joint i (i.e., the new virtual sensor which is denoted as Sv in Fig. 2) in joint frame i\u22121, i\u22121XS1 is the position of the real sensor located on link i\u2212 1 and i\u22121aS1 is its output. Given ia v is not affected by joint i rotation, it will be the linear acceleration of the i-th link after it is transformed into joint frame i, i.e. ial = ia v. iaSm in (3) is determined recursively from the first sensor installed on the first link to the last sensor located at the end-effector. One has to make sure that the correct pieces of information are sent between two adjacent links, i"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000586_0954409716631785-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000586_0954409716631785-Figure4-1.png",
        "caption": "Figure 4. Additional bending moments due to left (a) and right (b) flanging condition.",
        "texts": [
            " The element Balance2 is symmetrical to the first one, with the exception that instead of the longitudinal actuators two instrumented rods (Figure 3(d)) are used. The test rig allows only the static (non-rolling) calibration of the wheelset but, while operating, the wheelset can move laterally, thus the effects of the lateral displacements must be considered and taken into account during the calibration process. This is done by repeating the calibration process in three different relative lateral wheel\u2013rail positions: centred, left flanging and right flanging conditions. Indeed, as shown in Figure 4, it is possible to observe that, while flanging, additional bending moments in the vertical plane appear on the axle. The equations defining these bending moments can be easily derived and they are reported in equation (4). It is important to observe that these moments are mainly caused by the lateral displacement y of the actual contact point with respect to the nominal Table 1. Summary of the load condition applied in the calibration process. Load level X1 X2 (kN) 10 5 0 // // // // // at UNIV OF VIRGINIA on June 4, 2016pif"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure21.2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure21.2-1.png",
        "caption": "Fig. 21.2 Interaction scheme of man and the exoskeleton as the BTWS",
        "texts": [
            " At this stage, a BTWS providing an assistive movement of the operator\u2019s back is considered, due to the installation of an active hip joint (HJ) in the exoskeleton based on a combined LGC [4\u20136]. Such a biotechnological system, called BTWS, allows us to unload the lumbar spine muscle system, while the human-machine interface, which provides a control system for the interaction of the person and the exoskeleton, is characterized by relative simplicity, a minimal number of sensors, and the relative simplicity of the control algorithm. The human model consists of the thigh 3, the exoskeleton\u2019s back 4, and the operator\u2019s back 5. The movement of BTWS is considered in Fig. 21.2. There are two kinds of compensation systems, which can be distinguished: active, having a controlled electric drive and passive, without a controlled drive. Next, we dwell on a combined type of a linear gravity compensator (LGC), which uses the positive properties of passive and active LGCs [6]. The LGC scheme is shown in Fig. 21.3. LGC consists of four main elements: 1\u2014the rod, 2\u2014the body, 3\u2014the elastic element, 4\u2014the electric drive (not shown in the diagram). The combined LGC operation principle is that, the potential energy of the elastic element accumulates during its\u2019 compression, and then, when it is stretched, the energy is transmitted to the exoskeleton in the form of a moment MO43(\u03a61) applied by the HJ"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001329_ecce.2019.8912507-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001329_ecce.2019.8912507-Figure5-1.png",
        "caption": "Fig. 5: Control points for spatially triggered torque control",
        "texts": [
            " That is, small errors are ignored in order to maintain six-step, while larger errors will be corrected very fast and within the control resolution of eTs. In contrast to the flux, which is controlled at every sampling instant, the average torque is controlled at intervals of sixty electrical degrees. The reason for this reduced controller update rate is that the flux angle can only be manipulated at the corners of the hexagon, as shown in Fig. 4. Consequently, the torque feedback to the torque controller is the averaged torque over sixty degrees (see Fig. 5). The deadbeat torque control law (4) is obtained through inversion of the operating point model of the torque equation (3). T _ em  _ = 3 4 P         lpml _ R Ld cos( _  + (Ld \u2013 Lq)l _ 2 R LdLq cos(2 _  (3)  _ (k+1) \u2248         T _ em  _ -1 T _ em(k+1) \u2013  _ (k) where  _ (k+1) =  _ (k+1) +  _ (k) (4) In (4), the previous angle change  _ (k) needs to be subtracted, since it takes an additional spatial sample point until the torque change is reflected in the average torque. This is further explained in the example below in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001535_j.ymssp.2020.106723-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001535_j.ymssp.2020.106723-Figure7-1.png",
        "caption": "Fig. 7. Dynamic model of misaligned cardan shaft.",
        "texts": [
            " The cardan shaft installed in CRH train consists of two parts which allows for relative movement as shown in Fig. 6. One of the parts is steel spline shaft and the other part is steel hollow shaft. The parameters of the cardan shaft in CHR train are listed in Table 3. Table 2 Mean acceleration amplitude. Articulation angle antransport(m/s2) attransport(m/s2) anrelative(m/s2) aCoriolis(m/s2) 0 168.9 0 0 0 2 168.9 2.6 0.1 0.1 4 168.8 5.2 0.3 0.4 6 168.7 7.7 0.7 0.8 Fig. 6. Structure of CHR train cardan shaft. The diagram describing dynamics of misaligned cardan shaft in rotation is shown in Fig. 7. The forces generated including axial force and radial force. According to D\u2019 Alembert principle [28], the force is equal to mass times acceleration. The cardan shaft is approximately defined as a cylinder and the mass is distributed over an infinite number of circles whose center is the axis of the shaft. Assuming a mass element P on cardan shaft, the element P is on a plane which crosses point Q and perpendicular to line BC. The angle between QP and AB ish, and the distance between points P and Q is r"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002890_iccia52082.2021.9403563-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002890_iccia52082.2021.9403563-Figure1-1.png",
        "caption": "Fig. 1. Quadrotor geometric representation [18] II. PRELIMINARIES",
        "texts": [
            " Then, the Lyapunov function converges to zero in a finite time with the following settling time 1 1 1 1 0(1 ) ln(1 )T V \u03b3\u03b1 \u03b3 \u03b1\u03b2\u2212 \u2212 \u2212 \u2212\u2264 \u2212 + (4) Lemma 2 [20]: Consider the following fractional-integer differential equation: 1[ ( ) ] 0ax kD sig x\u03bb\u2212+ =& (5) where nx R\u2208 , ( )ik diag k= is a constant positive definite matrix with components ik R +\u2208 , 1,2,...,i n= and 1 1( ) [ ( ),..., ( )] a aa T n nsig x x sign x x sign x= with 0 , 1a \u03bb< < . Then, the state x converges to zero in a finite time named xT . In the next section, problem formulation including system description and the aim of this paper are presented. III. PROBLEM FORMULATION A quadcopter consists of four motors in cross situation as in Figure 1. These motors produce the necessary torque for attitude control. A quadcopter moves in three dimensions and consists of six degrees of freedom. However, given that it is mainly focused on the control of altitude and attitude of a quadrotor, its move can be described by a dynamic equation with four state variables. Considering [ , , , ]Tq z \u03c6 \u03b8 \u03c8= as the generalized coordinates in which z stands for altitude (height) and [ , , ]T\u03b7 \u03c6 \u03b8 \u03c8= denotes the attitude vector that its elements are respectively roll, pitch and yaw angle, the Lagrange dynamics of quadrotor Attitude/Altitude is as follows [18] 4 3 2 1 1 )( )( )( U II II U I l I J I II U I l I J I II g m U ccz zz yx yy r y xz xx r x zy + \u2212 = +\u2126+ \u2212 = +\u2126\u2212 \u2212 = \u2212= \u03b8\u03c6\u03c8 \u03c6\u03c8\u03c6\u03b8 \u03b8\u03c8\u03b8\u03c6 \u03b8\u03c6 &&&& &&&&& &&&&& && (6) where is defined as below [18]                           \u2212\u2212 \u2212 \u2212 =             2 4 2 3 2 2 2 1 4 3 2 1 00 00 \u03c9 \u03c9 \u03c9 \u03c9 dddd blbl blbl bbbb U U U U (7) where m stands for the mass of quadrotor, l is the distance from BO to the rotor center, , ,x y zI and 1,2,3,4\u03c9 are the quadrotor moments or inertia and angular velocities respectively and 1 2 3 4\u03c9 \u03c9 \u03c9 \u03c9\u2126 = \u2212 + \u2212 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002856_s13369-021-05654-z-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002856_s13369-021-05654-z-Figure6-1.png",
        "caption": "Fig. 6 Schematic of shot blasting experiment",
        "texts": [
            " Response surface methodology (RSM) with face-centered design (FCD) was employed in the present work to investigate the effect of input variables (viz., time, pressure and grit/particle size) on the surface roughness as the output variable. Input variables with their levels of experimentation are shown in Table\u00a03, and the resultant matrix involving 30 sets of experiments derived from RSM using Design Expert 11.0 software (Trial Version) is shown in Table\u00a04. Post-processing using shot blasting of the additively manufactured aluminum alloy specimens was carried out using shot blasting experimental setup, as shown schematically in Fig.\u00a06. In this experimental setup, the testing specimens were mounted perpendicular to the nozzle (dia. 5\u00a0mm) of the gun using a fixture, such that, the center axis of the gun\u2019s nozzle was concentric with the center of the rectangular workpiece. During the working condition, the shots of glass beads were continuously sucked through the grit suction inlet and sprayed with high air pressure onto the surface of testing specimen maintained at the stand-off distance of 80\u00a0mm. During this study, all the specimens were subjected to testing as per the design matrix shown in Table\u00a04, developed using DOE approach"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001217_j.mechmachtheory.2019.103633-Figure26-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001217_j.mechmachtheory.2019.103633-Figure26-1.png",
        "caption": "Fig. 26. The structural design model of ZSFH_4BSL.",
        "texts": [],
        "surrounding_texts": [
            "The sample of ZSFH_4BSL is made according to the ZSFH design parameters in Section 3.4.2 , and its 3D structure is shown in Figs. 26 and 27 shows a rectangular leaf spring string sample and a 4BSL sample. Fig. 28 shows a ZSFH_4BSL sample.\nAccording to the rotational stiffness test method proposed in [30] , the inner ring of ZSFH_4BSL is rotated by an angle through a precision angle measuring rotary table. At this time, the torque on the outer ring of ZSFH_4BSL is measured by the torque sensor and displayed on the sensor digital display meter. The test platform is shown in Fig. 29 .",
            "The rotational stiffness test was performed on the IORFH and ZSFH_4BSL. According to the Eqs. (27) and (25) , and the test results, the comparison chart of torque and rotation angle of IORFH and ZSFH_4BSL is shown in Fig. 30 . According to the Eq. (26) and the test results, the zero stiffness quality curve of ZSFH_4BSL is shown in Fig. 31 . The abnormal conditional of the sample during the experiment is shown in Fig. 32 .\nIt can be seen from Fig. 30 that the torque of ZSFH_4BSL is almost zero at a rotational angle of \u221220 \u00b0 to 20 \u00b0, and the difference between the theoretical values and the measured values is small. Therefore, the design purpose is achieved. In addition, the relationship of torque and rotation angle of IORFH as a positive stiffness module has high linearity. The error between the theoretical values and the measured values is small, which is consistent with the characteristics of IORFH."
        ]
    },
    {
        "image_filename": "designv11_14_0002542_icem49940.2020.9270965-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002542_icem49940.2020.9270965-Figure13-1.png",
        "caption": "Fig. 13. Structure of proposed ISNCW machines [28].",
        "texts": [
            " Adopted the AlNiCo PMs can be completely demagnetized in the post-fault mode and the motor will operate as switched reluctance motor, which Authorized licensed use limited to: University of Gothenburg. Downloaded on December 20,2020 at 19:19:58 UTC from IEEE Xplore. Restrictions apply. possessed inherently FT capability. Meanwhile, the novel topology has high power density and wide speed range simultaneously 3) Separated Phase Machine An integral slot non-overlapping concentrated winding (ISNCW) and machine configuration have been proposed [28], shown in Fig. 13. The topology has better inherently isolation capability between different phases and merits in lower short circuit current, FT capability. Furthermore, the separated phase winding structure have advantages in maintenance. In conclusion, conventional FT methods depends on isolation and SC current suppression have been summarized, as benchmark. Besides, some FT innovations applied on electric vehicle, which could pave the way for application on aircraft, have been introduced. Aforementioned dual three-phase are the common FT topology which given consideration to both technology readiness level and reliability"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003542_pime_proc_1948_158_045_02-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003542_pime_proc_1948_158_045_02-Figure5-1.png",
        "caption": "Fig. 5. Worn Profiles used on Model",
        "texts": [
            " They were provided with three interchangeable sets of flanges, having coning angles of 50 deg., 60 deg., and 70 deg. The reproduction of worn profiles one-tenth full size was difficult and expensive, and only a few could be made with the resources available. The first pair of worn tyres was ground so hollow that the effect of reversed coning was produced, and the axle ran hard over to either side. Two other pairs of worn tyre profiles, W1 and W2, and two of rail profiles, R1 and Rz, were made; they are shown in Fig. 5. One important effect of wear, the increase in the maximum difference between the effective radii, can be reproduced by the simple expedient of increasing the coning angle; and the authors felt that the use of this expedient might give the best indication of the general behaviour of worn wheels, as distinct from the idiosyncrasies of particular proiiles. Accordingly, a pair of wheels with treads coned at 1 in 5 was provided. With all wheels the flange clearance could be varied by inserting washers between the hubs and the shoulders of the axle",
            " the angles of incidence and reflection are about equal, the path is a sharp zig-zag from rail to rail, and the wavelength is consequently shortened; the deflexion of the rails considerably increases the amplitude of the motion. The records for the 1 in 5 coning show a shorter wavelength, in agreement with theory. The motion passes through much the same stages as the speed increases, but at any given speed it is much more violent; for instance, the motion at 20 ft. per sec. is very similar to that of the 1 in 15 coning at 35 ft. per sec. Earlier tests with 1 in 10 coning gave results intermediate between 1 in 5 and 1 in 15. Fig. 8 shows records for three combinations of the worn wheel and rail profiles of Fig. 5. The wavelengths are greater than with 1 in 5 coning, but considerably less than with 1 in 15. Each combination gave a distinctive path, the same tyre profile running differently on different rail profiles; some of the records in the middle are quite unlike those on their right and left, although they have one profile in common with each. This confirms that the running of worn wheels depends more on 0 25 50 75 100 INCHES Treads, 1 in 15. Treads, 1 in 5. Flanges, 60 deg. Flanges, 60 deg. Flange clearance, Flange clearance, 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001993_s00170-020-06278-7-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001993_s00170-020-06278-7-Figure8-1.png",
        "caption": "Fig. 8 Edge width according to contactFig. 7 Projected contact between the end disk wheel and the tool",
        "texts": [
            " Therefore, the minimum value ofD can meet the design requirements of the outer core radius with changing the value of \u03b1S. In this situation, the coordinates of point, B, can be calculated. XB YB \u00bc RE\u2219cos\u03b8B\u2219sin\u03b2 F \u00fe OWEx \u2212RE\u2219sin\u03b8B \u00fe OWEy \u00f07\u00de The above rule applies not only to the outer core radius but also to the second grinding flute end point, E, depending on the grinding path of the end disk wheel. Given the end disk wheel center point coordinates, [OWEx OWEy], the coordinates of point E can be calculated by dichotomy. In Fig. 8, the two corresponding angles of point E are \u03b8E and \u03b1E. ZE, the Z axis coordinate of the end disk wheel center, can be calculated with the angle \u03b8E. ZE \u00bc RE\u2219cos\u03b8E\u2219cos\u03b2 F \u00f08\u00de This decreases with the decrease in wheel width as the radius is ground by different disk wheels. There will always be one disk wheel forming the joint point radius, rj, as shown in Fig. 8. If the value of the wheel width changes by dichotomy, the radius at the joint point can be approximated. The method to calculate the coordinates of joint point, J, is the same as the one used for point B, so, it does not need to be further described. The corresponding angles at point J are represented by \u03b8J and \u03b1J. Given the angle, \u03b8J, the Z axis coordinate of the end disk wheel center, ZJ, can be calculated: Z J \u00bc RJ \u2219cos\u03b8 J \u2219cos\u03b2 F \u00fe hJ \u2219sin\u03b2 F \u00f09\u00de RJ \u00bc RE \u00fe hJ cot\u03c3 \u00f010\u00de where hJ is the width of the disk wheel that forms the joint point and RJ is the corresponding disk wheel radius",
            " The following coordinate transformation is therefore required: OWFx \u00bc OWEx \u00fe hw\u2219cos\u03b2 F OWFy \u00bc OWEy \u00f013\u00de For two-pass flute grinding, the grinding mark and edge width are determined by the value of the Z axis wheel center coordinate and the rotation angle of the tool. Figure 9 a shows the initial flute point, S, on the Y axis after the first grinding and the angle between the Y axis and point J, which is labeled Aff. However, the disk wheel that grinds the joint point for the second pass does not meet point J in this position. Figure 8 shows that the contact angle is \u03b1J. The tool therefore needs to rotate by an angle of ARO to ensure the accuracy of the second pass. The Z axis value of the end disk wheel center is ZJ when the wheel grinding point is J. The wheel center coordinates,OWFz, can then be described according to Eq. 15. Given the wheel center coordinates, OWFx OWFy OWFz\u00bd , and the rotation angle, ARO, the flute grinding at the second pass will combine with the first pass grinding without leaving a gear mark. ARO \u00bc Aff\u2212\u03b1 J \u00fe \u03c0=2 \u00f014\u00de OWFz \u00bc Z J \u00fe hw\u2219sin\u03b2 F \u00f015\u00de The complete computation process for two-pass flute grinding, based on the wheel center modes and wheel orientation, is given in this section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003720_s0039-9140(96)02040-1-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003720_s0039-9140(96)02040-1-Figure14-1.png",
        "caption": "Fig. 14. Microcell with working rotating disk electrode with impeller for batch injection analysis.",
        "texts": [
            " After the stripping stage the magnetic stirring is switched on again for cleaning of the working electrode. A microcell with a rotating working electrode [5] has been used for batch injection voltammetric analysis. In this cell (Fig. 13), (1) is a large compartment, (2) is the large-volume supporting electrolyte, (3) is the rotating disk working electrode (glassy carbon), (4) is the auxiliary electrode, (5) is the reference electrode and (6) is the plastic tip for sample injection by micropipette. A cell [99] with a rotating disk working electrode (Fig. 14) was used for macrosamples but it can also be applied for batch injection voltammetric microanalysis. In Fig. 14, (1) is the compartment, (2) is the supporting electrolyte, (3) is the rotating disk working electrode (glassy carbon), (4) is the impeller fixed to the shaft of the rotating electrode and rotating together with it, (5) is the auxiliary electrode, (6) is the reference electrode, (7) is the channel for internal circulation of electrolyte, and (8) is the micropipette for inserting the sample in channel (7). In the case of a working rotating disk electrode with an impeller, additional internal circulation of electrolyte occurs. The inlet sample is first moved down in channel (7) and then up to the working electrode surface. On contact of the sample with the electrode surface the analytical signal is generated. Such a system is analogous to the usual system for detection in flow injection analysis. However, the system [99] is different from the others in its simplicity and compactness. In comparison with the microcell in Fig. 13, in the microcell in Fig. 14 steadier movement of the sample to the working electrode surface and a more intensive hydrodynamic effect have been achieved [99]. We have seen that the microcell with a static working electrode (Fig. 12) was used for batch injection in anodic stripping voltammetry and this design did not allow the use of a rotating working electrode for stripping voltammetric analysis. At the same time, the design in Fig. 15 [36, 37] with the rotating disk working electrode in the microcell for batch injection in anodic stripping voltammetry decreased the detection limit by approximately an order of magnitude"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001287_8756087919887216-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001287_8756087919887216-Figure2-1.png",
        "caption": "Figure 2. Mathematical illustration for the region under study.",
        "texts": [
            " The experimental relations of the Rabinowitsch model for two-dimensional fluid flows are sxx \u00bc 2l @ u @ x sxy \u00fe K\u00f0 sxy\u00de3 \u00bc 2l @ u @ y \u00fe @ v @ x syy \u00bc 2l @ v @ y 9>>>>= >>>>; (8) where The suitable initial and boundary conditions corresponding to the flow problem are sxy \u00bc 0 at y \u00bc 0 u \u00bc U at y \u00bc r\u00f0 x\u00de (9) For non-dimensional expressions, we introduce variables defined by With the help of above defined expressions, equations (5) to (9) become @u @x \u00fe @v @y \u00bc 0 (10) bRe u @u @x \u00fe v @u @y \u00bc @p @x \u00fe b2 @sxx @x \u00fe b @sxy @y (11) l \u00bc zero-shear viscosity K \u00bc coefficient of pseudoplasticity x \u00bc xffiffiffiffiffiffiffi RH0 p axis in the flow direction u\u00f0x; y\u00de \u00bc u\u00f0 x ; y\u00de U velocity of the fluid in the x-direction v\u00f0x; y\u00de \u00bc v\u00f0 x; y\u00de U ffiffiffiffi R H0 q fluid velocity in the y-direction p \u00bc H0 p lU ffiffiffiffi R H0 q pressure sxy \u00bc H0 sxy lU shear stress Q \u00bc Q 2UH0 flow rate a \u00bc Kl2U2 H0 2 Rabinowitsch parameter r \u00bc h\u00f0x\u00de H0 \u00bc 1\u00fe x2 2 dimensionless width from the center line to the roll surface k \u00bc H H0 coating thickness b2Re u @v @x \u00fe v @v @y \u00bc @p @y \u00fe b2 @sxy @x \u00fe b @syy @y (12) sxx \u00bc 2b @u @x sxy \u00fe a\u00f0sxy\u00de3 \u00bc @u @y \u00fe b2 @v @x syy \u00bc 2b @v @y 9>>>= >>>; (13) Here, b \u00bc ffiffiffiffi H0 R q is the geometric parameter and Re \u00bc qUH0 l is the Reynold\u2019s number. where The corresponding boundary conditions are sxy \u00bc 0 at y \u00bc 0 u \u00bc 1 at y \u00bc r\u00f0x\u00de (14) Moreover, the boundary condition at the symmetric line of u at y \u00bc 0 is @u @y \u00bc 0when y \u00bc 0 (15) At the symmetric line, the separation point xs needs to be calculated and two additional boundary conditions are required to calculate its numeric value. From the physics of the problem, the stagnation point is at the axis of symmetry as shown in Figure 2, and one can define the following boundary conditions u \u00bc 0 when x \u00bc xs and y \u00bc 0 (16) The boundary condition at the separation point xs for the pressure or force equates the pressure p to the pressure related with surface tension c. p \u00bc c r atx \u00bc xs p \u00bc H0 r \u00f0Nca2\u00de 1 at x \u00bc xs 9>= >; (17) The parameter NCa2 \u00bc lU c R H0 1 2 is a modified capillary number. P \u00bc density U \u00bc velocity H0 \u00bc nip separation l \u00bc zero-shear viscosity Figure 2 indicates that a good approximation to the free surface is 2r\u00fe 2H \u00bc 2H0r\u00f0xs\u00de r H0 \u00bc rs k 9= ; (18) where k \u00bc H H0 is a non-dimensional parameter for coating thickness and will be the key factor that we seek to predict in this model. Furthermore, we know that r \u00bc 1\u00fe 1 2 x 2, and from this one can define rs \u00bc r\u00f0xs\u00de \u00bc 1\u00fe 1 2 xs 2 (19) In a roll coating analysis, both Re and b are much smaller than unity. Under these assumptions, equations (11) to (13) reduce to @p @x \u00bc @sxy @y @p @y \u00bc 0 9>= >; (20) sxx \u00bc 0 sxy \u00fe a\u00f0sxy\u00de3 \u00bc @u @y syy \u00bc 0 9>= >; (21) In equation (20), it is clear that p is not a function of y and is function of x only",
            " Numerical solution and engineering parameters of the process The governing boundary value problem can be solved exactly, therefore integrating equation (20) and using boundary condition (14), one can write that sxy \u00bc dp dx y (22) Using equation (22) into equation (21), we obtain @u @y \u00bc y dp dx \u00fe ay3 dp dx 3 (23) Integrating again and using the remaining boundary conditions (14) gives u\u00f0x; y\u00de \u00bc 1 2 dp dx \u00f0y2 r2\u00de \u00fe a 4 dp dx 3 \u00f0y4 r4\u00de \u00fe 1 (24) The above equation contains dp dx, which is still unknown and can be determined by introducing the dimensionless volumetric flow rate as Q \u00bc k \u00bc Z r 0 udy (25) Inserting equation (24) into equation (25) and subsequent integration yields ar5 5 dp dx 3 \u00fe r3 3 dp dx \u00fe \u00f0k r\u00de \u00bc 0 (26) The above equation is cubic in dp dx and its real solution is dp dx \u00bc 5 4 1 3 2 3 5 1 3ar5 \u00fe 2 1 3 27a2r10\u00f0k r\u00de \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 20a3r24 \u00fe 729r20\u00f0k r\u00de2a4 q 27a2r10\u00f0k r\u00de \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 20a3r24 \u00fe 729r20\u00f0k r\u00de2a4 q ar5 0 BB@ 1 CCA (27) Since r\u00f0x\u00de is an algebraic function, one may integrate equation (27) to get the pressure profile p\u00f0x\u00de as p\u00f0x\u00de \u00bc 5 4 1 3 Z x 1 2 3 5 1 3ar5 \u00fe 2 1 3 27a2r10\u00f0k r\u00de \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 20a3r24 \u00fe 729r20\u00f0k r\u00de2a4 q 27a2r10\u00f0k r\u00de \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 20a3r24 \u00fe 729r20\u00f0k r\u00de2a4 q ar5 0 BB@ 1 CCAdx (28) When equation (17) is used in above equation, an integral equation in k is obtained as 1 \u00f0rs k\u00deNca2 \u00bc 5 4 1 3 Z xx 1 2 3 5 1 3ar5 \u00fe 2 1 3 27a2r10\u00f0k r\u00de \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 20a3r24 \u00fe 729r20\u00f0k r\u00de2a4 q 27a2r10\u00f0k r\u00de \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 20a3r24 \u00fe 729r20\u00f0k r\u00de2a4 q ar5 0 BB@ 1 CCAdx (29) As yet separation point xs is unknown. For this, substitute Eq. (27) into equation (24) and use boundary equation (16) at the separation point, as an explicit relation between the dimensionless flow rate k, and the separation point xs is needed. In case of the constitutive model under study, it is not possible to find such a relationship. Figure 3 shows the complex implicit relation between separation point and coating thickness. It is quite evident from Figure 2 that the film splits uniformly; therefore, the separation point is \u00f0xs; 0\u00de. Velocity and pressure are zero at this position. From the implicit relation between separation point xs and dimensionless coating thickness k, one can easily generate the data in order to interpolate the desire polynomial. Table 1 is the generated data with a \u00bc 50. A 10 degree interpolating polynomial has been constructed in terms of coating thickness, which has been further used in equation (29) in order to find the desired coating thickness"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002671_j.jsv.2021.115967-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002671_j.jsv.2021.115967-Figure9-1.png",
        "caption": "Fig. 9. The numerical model for movement in the yz-plane (a) the two degree-of-freedom system model, (b) first mode shape, and (c) second mode shape (corresponding to the fourth experimental mode).",
        "texts": [
            "2 s after each impact) in a matrix, and we estimate the modal parameters from this matrix using a stabilization diagram based on the Ibrahim timedomain method [2,20] , see Fig. 7 . Fig. 8 illustrates the five mode shapes from the identification process and Table 1 holds the identified modal parameters. We create a simple finite element model of the test-specimen simplified to 2D - corresponding to the yz-plane from Fig. 6 - with one translational and one rotational DOF using an Euler-Bernoulli beam element fixed at one end and with a lumped mass with mass moment of inertia in the other end, see the system in Fig. 9 (a). The first and fourth mode from the identification process from the experimental analysis correspond to the given plane. The model is updated to resemble the experimental analysis in terms of natural frequencies, see the modal parameter in Table 1 . Fig. 9 (b-c) illustrates the two mode shapes of the system. We use Eq. (18) and the mode shapes from Table 1 to plot the frequency response function matrix of the measured acceleration in Fig. 10 for both setups. We apply an impact load next to and in the direction of one of the lasers and measure both displacement from the lasers and the acceleration from the accelerometers. This results in a free decay of the test specimen in the yz-plane. Assuming for the given frequency range of interest, the wooden block is rigid and we calculate the acceleration of the block as a mean value",
            " The amplitude difference - based on the analytic expression for tilt - corresponds well with the ratio of standard deviation between accelerometers and lasers. We apply these amplitude differences to adjust the measured acceleration from the accelerometer, see Fig. 12 . This corrected the amplitude of the acceleration from the accelerometers so they have similar amplitude as the lasers. Similarly, we adjust the estimated mode shapes for the tilt effect in Fig. 13 where we obtain a better resemblance to the mode shapes from the model, see Fig. 9 . Here the MAC-value increases to above 0.99 for the first mode shape as compared to the values in Table 1 . Finally, we estimate the angular displacement by the method from from section 4.1 using the four vertical sensors (unaffected by translational motion) and we conclude that the assumption of small angles is valid for the given case, see Fig. 14 . This estimation of angular displacement only requires the vertical displacement and the geometrical distances to the centre of rotation. This tilt sensing application introduces a small amount of noise in the estimated angular displacement so they do not form a free decay of a single mode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001692_012041-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001692_012041-Figure5-1.png",
        "caption": "Figure 5. Bionic substantiation of the model of ring-cutting working bodies of the ice rink: a - the digging limb of the common rhinoceros beetle (Orictes nasicornis); b is a side view of the model of ring-cutting working bodies of the roller.",
        "texts": [
            " The values shown in Figure 4 (a) and (b) were calculated using a Lenovo ideapad 310-15 IAP - 1 laptop, ZET 017- T8 - 2 strain gage, ZET017-U2 - 3 analyzer, TS21-T2 - 4 strain gage and two piezoelectric accelerometers BC 110 - 5. The object of research is the technological process of interaction of ring-cutting working bodies of the roller with soil. As applied to the loosening working bodies of the ice rink, exploratory studies have established that the common rhinoceros beetle (Orictes nasicornis) performs efficient movement in the soil with front paws with teeth of b width and located with a certain step S, one of which is shown in Figure 5 (a). The ring-cutting roller model proposed in Figure 5 (b) also contains ripping elements 2 located with a certain step S on the rotary disks 3 [11]. The ripping elements 2 of the rotary discs 3 of the ring-cutting roller model 1 have the shape of truncated cones, the lateral surface of which is made along the length of the logarithmic spiral f (x), as can be seen in Figure 5 (b). Experimental verification of traction resistance of the ring-cutting roller model shown in Figure 6 (a), in comparison with the 3KKSh-6 serial roller, performed in the experimental field of V.I. Vernadsky of Crimean Federal University, confirmed the feasibility of its use. During field trials, soil moisture was 15 ... 17%, soil hardness \u2013 38.9 ... 51.4 N / cm2, deformation index \u2013 2.8 \u2022 10-7 ... 3.5 \u2022 10-7 m2 / N, soil type - southern black earth. Figure 6 (b) shows the graphical dependence of the traction resistance of the ring-cutting roller model on the speed of movement in comparison with the serial model 3RSR-6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002085_tec.2020.3044000-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002085_tec.2020.3044000-Figure12-1.png",
        "caption": "Fig. 12. Electromagnetic forces acting on the rotor for the 39-slot with 12-pole motor",
        "texts": [
            " It can be generalized that if the number of slots is increased in asymmetric and unbalanced winding motors, less UMP force components are generated. For low number of slots and poles UMP force component would be significant. All the asymmetric and unbalanced structures investigated in this paper have UMP forces. The key issue is that the less the asymmetry is, the lower the UMP force component is. For instance, the winding arrangement of 39-slot with 12-pole machine used as a case study shown in Fig. 4 is also distributed asymmetrically producing UMP force component (Fig. 12). FEA simulations are carried out under rated current and electromagnetic forces acting on the rotor for different rotor positions are obtained (Fig. 13). An average force of 10.4 N is arising on the rotor. The variation of UMP force components is not centered on the origin due to the asymmetrical disposition of the phase windings. However, it Authorized licensed use limited to: University of Gothenburg. Downloaded on December 18,2020 at 20:38:30 UTC from IEEE Xplore. Restrictions apply. 0885-8969 (c) 2020 IEEE"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003907_nme.1620380611-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003907_nme.1620380611-Figure4-1.png",
        "caption": "Figure 4. Helicoidal staircase with both ends fixed",
        "texts": [
            " The results obtained from the model of this study for the ring (or for the long cylindrical pipe) under the single load, P, in its plane are shown in Table I. These results are generally in close agreement with those of the literature.6 The author believes that the numerical values given in the literature for U, and M b for the cases: 4 = NO\", 4 = 45\" and 4 = 315\", respectively, which are rewritten here in Table I, are incorrect probably due to misprinting. Example ZI: Helicoidal staircase with constant cross-section. This system is subjected to uniformly distributed vertical loading (Figure 4). The distance between the resultant of the distributed load and the axis of the rod is defined as eccentricity, which is equal to e = b2/(12a). This eccentricity always causes an extra distributed moment. The results obtained for this example by Ta bl e I. R es ul ts f or th e ex am pl e pr ob le m o f a rin g el as tic al ly s up po rt ed in th e ra di al d ire ct io n (F ig ur e 3) ut (m ) U \" (m ) fi b (r ad ) { TI (N ) Tn (N ) M b (N m ) { 1 0- 2) d (O ) R ef 6 Re f. 1 C FM R ef 6 C FM Re f",
            " Additionally, neglecting the eccentricity results in a difference of 50 per cent in Q,, 140 per cent in M , , and 42 per cent in M,. Therefore, especially for wide and shallow sections, the effect of eccentricity of the resultant load should be taken into account. Example 111. Helicoidal staircase with both ends $xed and having two intermediate universal joint supports under uniformly distributed vertical loading. The dimensions and material properties of the constant cross-section, geometric properties of the circular helix and distributed loads are given in Figure 4. The total angle of the system is 371. At 4 = 71 and 4 = 271 sections, there exist two universal joint supports, each of which prevents only the translational displacement at that support. The results obtained for this example are given in Table IV. Example I V: Helicoidal bar with a linear variation of the circular cross-section. The boundary conditions, the loading and the material properties of the helix are as same as in the second example above expect the cross-section is circular. At the fixed ends the radius of the cross-section is ro = 15 cm and at the midspan rl "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000212_mees.2019.8896501-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000212_mees.2019.8896501-Figure7-1.png",
        "caption": "Fig. 7. Results of the analysis of the radial AMB: a) distribution of the magnetic field lines (x, y=0 and x, y\u2248\u03b4r/cos45\u00ba); b) the power characteristic (a dependence of the total force modulus on the displacement).",
        "texts": [
            " They are indicated by dots in Fig. 6. The total of such calculations is about (2n+1)2. A static electromagnetic analysis is performed and the magnetic forces in the x, y directions and the total vector are determined at this mutual position (Fig. 6b). The results of analyzes at different rotor positions for a radial AMB with an outer stator diameter of 0.1 m and an internal pole diameter of 0.06 m with a nominal radial clearance of r=1 mm, taking into account the control law, are shown in Fig. 6b and 7\u0430. Fig. 7b shows the three-dimensional stiffness characteristic. It is obtained by approximating the magnetic force values at calculated points by complete polynomials of two variables x and y, followed by a differentiation. The calculation results for the radial AMB (Fig. 6, 7), as well as for the axial AMB (Fig. 2, 3), indicate that a negative feedback in the CS (see Fig. 4) and the selected control law parameters ensure a centering of the rotor in the gap. IV. CONCLUSION AND DISCUSSION The paper proposes numerical finite element methods for determining the force and stiffness characteristics of AMBs for almost any configuration and design"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001468_iccas47443.2019.8971725-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001468_iccas47443.2019.8971725-Figure7-1.png",
        "caption": "Fig. 7. Phasor diagram of the interferometric signals with the measured data phasors in red, the target phasors in blue and the clutter phasors in green. In (a), the true case and in (b), the case in which it is assumed \u03d5A = 0.",
        "texts": [
            " Since we simulated the ROC for a minimum PFA of 10\u22124 we used in all the simulations a sample size of 105. We note that, in this case, the PD will be affected by the estimation of the three parameters, thus, differently from the LRT case, it now depends also on the estimation of the target phase \u03d5A. The actual value of \u03d5A does not influence the velocity estimation accuracy and the GLRT performance, provided that it is properly estimated. For better evidencing the importance of estimating \u03d5A, we can refer to the phasor representation shown in Fig. 7. In Fig. 7(a), the phasor diagram of the interferometric signals for the case N = 2 is shown. The phasors of the measured data Z1 and Z2 are represented in red. They are resulting from the sum of the target phasors (in blue) and the clutter phasors (in green). The target phasors exhibit the same amplitude A, the same initial phase offset \u03d5A, and different phases, due to the radial velocity phase term \u03d5v . The clutter phasors, instead, due to the very high clutter correlation, are not changing from one acquisition to the other (thermal noise has not been considered, as its effect is equivalent to the production of a negligible change of the clutter phasors that can be neglected for high CNR values). The correct estimation procedure essentially consists in searching the target phasors with equal amplitudes, different phases and the same initial phase offset, together with the fully correlated clutter phasors, providing a sum equal to the data phasors. The phasors that would be obtained from an estimation procedure not involving the estimation of the phase offset \u03d5A and assuming a realvalued target (\u03d5A = 0) are shown in Fig. 7(b). As it is evident, the solution obtained by considering \u03d5A = 0 is quite different from the real one and leads to a wrong estimate of the phase related to the radial velocity. Anyway, for very high SCR values, when clutter can be neglected respect to the target signal, the velocity estimate becomes independent from the phase \u03d5A, as expected. In Fig. 8, we report the comparison between the LRT (point dashed line) and the GLRT (solid line) proposed ATI detectors on simulated data using TSX parameters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003484_9781119711230-Figure12.2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003484_9781119711230-Figure12.2-1.png",
        "caption": "Figure 12.2 An example of a fixed-wing drone.",
        "texts": [
            " \u2022 High-altitude platforms: These quasi-stationary platforms are used in some available applications for wide coverage, long endurance, and altitudes of about 17 km [8]. \u2022 Low-altitude platforms: These platforms are available in some available applications for quick mobility, fast and flexible deployment, and are cost-effective and typically file up to several hours. 2) Type: This is the second classification of the UAVs [10]. There are two categories of UAVs \u2013 fixed-wing drones and rotarywing drones. \u2022 Fixed-wing drones: These UAVs are renowned for military and resistance related applications. As shown in Figure 12.2, they simply utilize a wing like a normal airplane to fly. They do not need to exert a lot of energy to stay afloat in the air [11]. Their rigid structure generates lift under the wing due to forward airspeed. Because they use less energy they can fly for a longer time, and are utilized to monitor areas and survey focal points on the ground [11]. They use gas motors rather than electric motors, which increases their flying limit for as long as 16 hours [2]. The disadvantage of these kinds of automatons is that they can\u2019t maintain drifting positions, which makes them futile for photography, and they are hard to proper and land"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002900_j.isatra.2021.04.043-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002900_j.isatra.2021.04.043-Figure3-1.png",
        "caption": "Fig. 3. The tilt-rotor CUAV aircraft design used in this paper.",
        "texts": [
            " Paper\u2019s organization Section 2 describes the CUAV\u2019s nonlinear dynamics with aeroynamics; besides, the problem statement is formally given. The ain result is composed of the control algorithm for the CUAV ogether with the control allocation problem; these are given in ection 3. Section 4 presents simulations, including several cases nd comparisons. Real flight experiments and comparisons with orks that present experiments are given in Section 5. Finally, ome conclusions are discussed in Section 6. . System description and problem statement This section introduces the longitudinal mathematical model f the CUAV depicted in Fig. 3. Besides, we develop the main deas behind the transition maneuver control with the aim of resenting the problem formulation. 2.1. CUAV\u2019s mathematical model The CUAV can fly in three different operating conditions depending on its four motors\u2019 angular position, as is depicted in Fig. 1. The motors\u2019 angular position is represented by \u03b3 in Fig. 4. To achieve a transition (and back transition maneuver), the CUAV\u2019s rotors must simultaneously tilt from 0\u25e6 to 90\u25e6 (and from 90\u25e6 to 0\u25e6). This results in three flight modes described as follows: 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001312_tie.2019.2952780-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001312_tie.2019.2952780-Figure2-1.png",
        "caption": "Fig. 2. Colenoid type electromagnetic linear actuator: (a) 3-D schematic diagram of the electromagnetic linear actuator highlighting different parts, (b) comparison of the clamping force with the variation of the current in the coils for the simulated and manufactured models [13].",
        "texts": [
            " [14] proposed the design of an axial-gap helical-motion PM motor based thrust actuator which is capable of producing around 2KN of thrust, however the design such an actuator system is quite complicated considering their structural design concept. A survey of the available literatures shows that the force generation capability of the proposed electromagnetic actuators is not in kN range, which is a basic requirement of the lathe machine actuator to ensure a thrust force of 50 kN. Recently, to substitute the hydraulic system, as shown in Fig. 2, a colenoid (COL) type electromagnetic linear thrust actuator was developed by D. Han et al [15] which can exert a high clamping force of kN range. As per the simulation and the experimental data mentioned in [15], the actuator can ensure a force below 40 kN. It is noteworthy that the design employs an airgap of 0.2 mm between the stator and mover to ensure such a force level. The performance of the actuator is highly dependent on the variation of the airgap and the airgap can vary based on the elastic deformation of the stator-mover sections"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001485_j.mechmachtheory.2019.103776-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001485_j.mechmachtheory.2019.103776-Figure1-1.png",
        "caption": "Fig. 1. (a) A two-loop linkage, (b) its directed topological graph [10] .",
        "texts": [
            " The other is the sequence matrix which shows the relative sequences of each joint in corresponding FC. The sequence matrix is newly introduced in this paper and will be used to construct matrix expressions of second-order kinematic constraints. 2.1. Circuit matrix of fundamental cycles (FCs) Topological graph ( B, J ) is introduced for multi-loop linkages [19] . The vertex set B represents the bodies/links, and the edge set J represents the joints. All bodies in B and joints in J are indexed with natural numbers. A two-loop linkage is shown in Fig. 1 (a), and the links are indexed with numbers in circles. Spanning trees and co-trees are used to identify the FCs once the topological graph ( B, J ) is introduced. A FC is a closed-path containing exactly one edge of the co-trees. The number of FCs of a connected, planar graph is determined by the Euler number \u03b3 : = | J |-| B | + 1, in which | J |: = n is the number of joints, | B | is the number of bodies. The k th FC is denoted as k ( k = 1 , \u00b7 \u00b7 \u00b7 , \u03b3 ) , and the i th joint belonging to the k th FC is indicated by J i \u2208 k . A directed graph is introduced with arbitrarily assigned directions of edges, which represent the orientations of relative motions of two adjacent links. The orientation of a FC is defined that it is equal to the orientation of its co-tree, and is used to indicate the relative orientations of edges in corresponding FC. Because spanning trees and co-trees are not unique, the FCs are not unique either. The directed graph of the linkage in Fig. 1 (a) with two-FCs are shown in Fig. 1 (b). The orientations of 1 and 2 are defined by the orientations of co-tree edges J 4 and J 6 . The incidence and orientations of joints relative to each FC are indicated by the circuit matrix , denoted by B . There are \u03b3 rows corresponding to the \u03b3 FCs, and the n columns corresponding to the n joints of multi-loop linkages. The entries of circuit matrix B is defined as B ( k, j ) := \u23a7 \u23aa \u23a8 \u23aa \u23a9 1 , if J j is aligned with k , \u22121 , if J j is not aligned with k , ( k = 1 , 2 , \u00b7 \u00b7 \u00b7 , \u03b3 j = 1 , 2 , \u00b7 \u00b7 \u00b7 , n ) 0 , J j / \u2208 k , . (1) Thus, the circuit matrix of the directed topological graph of the linkage in Fig. 1 (b) is J 1 J 2 J 3 J 4 J 5 J 6 B = ( 1 1 1 1 0 0 0 1 1 0 1 1 ) 1 2 . (2) 2.2. Sequence matrix of fundamental cycles (FCs) The circuit matrices are sufficient for constructing matrix expressions of first-order kinematic constraints of multi-loop linkages with screw coordinates but not for the second-order ones, in which Lie product of screws are involved, and the unique predecessor of each joint needs to be determined firstly. To this end, one link or edge in each FC is chosen to be cut open virtually, which gives rise to a tree topology",
            " < k i j , ( i 1 , i 2 , \u00b7 \u00b7 \u00b7 , i j \u2208 k ) (3) in which i 1 denotes the root joint, and i j denotes the terminal joint. These ordering sequences can be expressed in matrix form as follows, and is termed as the sequence matrix . The sequence matrix of the k th FC, k ( k = 1 , \u00b7 \u00b7 \u00b7 , \u03b3 ) , is denoted as k , with entries defined as, k ( i, j ) := \u23a7 \u23a8 \u23a9 1 , \u22121 , 0 , if i < k j, if i > k j, if i = j , or either J i or J j / \u2208 k , ( k = 1 , 2 , \u00b7 \u00b7 \u00b7 , \u03b3 i, j = 1 , 2 , \u00b7 \u00b7 \u00b7 , n ) . (4) For the graph of the linkage with two FCs in Fig. 1 (b), the two ordering sequences are 1 : 1 < 1 2 < 1 3 < 1 4 , 2 : 2 < 2 3 < 2 5 < 2 6 , (5) in which the joint J 1 and J 2 are the root joints of the respective FCs. The sequence matrices corresponding to 1 and 2 are given by Eq. (4) as J 1 J 2 J 3 J 4 J 5 J 6 1 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 1 1 1 0 0 \u22121 0 1 1 0 0 \u22121 \u22121 0 1 0 0 \u22121 \u22121 \u22121 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 J 1 J 2 J 3 J 4 J 5 J 6 , J 1 J 2 J 3 J 4 J 5 J 6 2 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 0 0 0 0 0 0 0 1 0 1 1 0 \u22121 0 0 1 1 0 0 0 0 0 0 0 \u22121 \u22121 0 0 1 0 \u22121 \u22121 0 \u22121 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 J 1 J 2 J 3 J 4 J 5 J 6 , (6) which are obviously skew-symmetric"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002037_s00170-020-06366-8-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002037_s00170-020-06366-8-Figure1-1.png",
        "caption": "Fig. 1 Schematic diagram of PBF-MSHAM strategy",
        "texts": [
            " In the first method, the cost to produce the precision assembly alignment features for the pre-machined body and the AM insert portion could be reasonably high. In the second, the additive process might fail, thus necessitating the re-preparation of the complex pre-machined near-net-shape body. One possible way to overcome these potential flaws is to use a very simple pre-machined part as the base of the additive build process; a strategy referred to as PBFmachined substrate hybrid additive manufacturing (PBFMSHAM). Figure 1 shows the schematic diagram of the proposed PBF-MSHAM strategy. In this study, PBF-MSHASM is defined as the serial process in which substrate parts are firstly prepared by conventional machining, then aligned and mounted onto the build plate of a powder bed fusion system. The powder is then filled up to the top level of the substrate part, and the AM part is then built directly onto the substrate part. After the build process, the part is removed from the build plate and the integrated base and AM part is finish-machined to achieve the required tolerances and surface finish"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001483_s42835-020-00365-1-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001483_s42835-020-00365-1-Figure2-1.png",
        "caption": "Fig. 2 Structure of GPS\u2013GLONASS antenna",
        "texts": [
            " Section\u00a02 demonstrates the equivalent circuit analysis of square patch antenna to find the RLC values using MATLAB program. Section\u00a03 gives an account of proposed 3D antenna design using CST software tool in order to find the return loss, gain, voltage standing wave ratio (VSWR) and radiation pattern of the antenna. Section\u00a0 4 illustrates the simulation results and Sect.\u00a0 5 describes the prototype of the antenna with the measured results. Finally, the Sect.\u00a06 spotlights the return loss S11 along with the conclusion and scopes of works in navigation system. The proposed antenna geometry is shown in Fig.\u00a02, where a coaxial-fed simple square patch circularly polarized antenna is used to combine the frequency of GPS and GLONASS systems. Figure\u00a03 shows the equivalent circuit for the square patch antenna. The truncation at the corners of the square patch is to obtain the circular polarization of the antenna. The Coaxial feed dimension is calculated to match 50\u00a0\u2126 impedance. The dimensions of the antenna are optimized using Computer Simulation Technique (CST) software for dual frequency operation and it is the fast and accurate 3D simulation of high frequency devices"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002221_0278364919897134-Figure18-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002221_0278364919897134-Figure18-1.png",
        "caption": "Fig. 18. Comparison between modeled and experimental fiber reorientation for the spherical membrane (top) and the conical membrane (bottom). Inflated volume increases for both samples from left to right. The theoretical fiber paths are superimposed upon images of the membranes from the inflation test. Note that the largest errors are in (d), where the spherical membrane approaches its critical volume.",
        "texts": [
            " As expected, the radial fibers straightened throughout the inflation process for both fiber-reinforced membranes. While the circular fibers in both samples successfully prevented radial expansion, the membranes did not reach their fiber-imposed volume limits owing to the stiffness of the elastomer, resulting in the curved portions of fiber at maximum inflation, especially in the conical case. The error between the theoretically-calculated and experimentally-observed fiber patterns can be found in Table 1 and Figure 18. With the exception of the last, nearcritical-volume test case of the spherical membrane, the model and experiment show good agreement with an average R2 value of 0.982 (62:598) for the spherical membrane and 0.995 (61:568) for the conical membrane. Here, we report the results of our internal and external load tests on spherical fiber-reinforced and unreinforced membranes. 5.3.1. External loading. Both the unreinforced and the spherical fiber-reinforced membranes were capable of lifting and supporting the test mass at 11",
            " While the wider pressure range used for inflation of a fiber-reinforced membrane has its advantages, it also makes actuating the membrane to its critical design case without failure of either the membrane or the clamp much more difficult. Owing to the geometry of the fiber layups designed for this article, the elastomer towards the edges of the membrane must deform significantly more than the elastomer in the center of the sample. This results in unmodeled deviations from the predicted fiber pattern close to the critical case (e.g., Figure 18(d)) and increased stress and deformation at the clamping surface. These stress concentrations lead to failure of the conical membranes before they can approach the critical case, but they can be mitigated in the future either by choosing a softer elastomer matrix or by modeling the stress\u2013strain relationship within the membrane to design fiber patterns that minimize stress concentrations. As shown in Table 1, the volume recorded at 0% inflation pressure is 36 cm3. This is due to the clamping force that is exerted around the edge of the membrane to maintain a seal against the base"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003681_971510-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003681_971510-Figure4-1.png",
        "caption": "Figure 4. Finite element model of anchor braket",
        "texts": [
            " COMPUTATION PROCEDURE 'The finite element moldel of caliper consists of 2130 solid elements and 3294 nodes as shown in Figure 2: At each time step, the computation starts with solving the dynamic equations. of motion to obtained the displacements, velocities and accelerations at each node by using Equations 1 through 4. The variation of contact surfaces are then determined by checking and computing the penetrations The finite element model of pads and backing plates consist of 756 solid elements and 1182 nodes as shown in Figure 3: The finite element of anchor braket consists of 639 solid elements and 1224 nodes as shown in Figure 4: The brake system consists of 5390 elements and 9368 nodes. FRICTION CURVES The coefficient of friction used in this case is a function of contact pressure and sliding velocity. Since the occurrence of brake squeal can be detected in a very short period of time, the variation of temperature has not been taken into consideration in this example. The friction data generated from the dynamometer test is shown in Figure 7. The pistons and fluid are modeled by using rods and springs respectively, as shown in Figure 5: RESULTS PISTON FORCES The time history of the force applied to each piston is shown in Figure 8"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001993_s00170-020-06278-7-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001993_s00170-020-06278-7-Figure6-1.png",
        "caption": "Fig. 6 End disk wheel grinding path",
        "texts": [
            " The value of the given wheel width hw should be larger than hmin in the first pass grinding calculation. With its given value, hw, the end disk radius of the 1V1 wheel can be calculated (see Fig. 5). RE \u00bc R\u2212hW cot\u03c3 \u00f02\u00de The radius at any position on the grinding wheel, RE, is described by Eq. (2), where R is the front disk radius of the wheel; and \u03c3 is the inclined angle. A 1V1 wheel can be regarded as being composed of a finite number of thin disks of different radii. Each disk generates one grinding path and these grinding paths form the flute envelope line, as shown in Fig. 6. The envelop line of all grinding paths that wheel disks generate forms the flute curve. However, except for grinding path that generates by end disk of grinding wheel, the paths generated by other disks cannot determine the externally tangent circle radius directly. These need a reference, and the end disk wheel can be the reference because the externally tangent circle radius of end disk wheel grinding path is outer core radius Rcb. During the flute grinding process, the wheel moves along the axis of the tool and the tool rotates anticlockwise"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001500_iccma46720.2019.8988753-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001500_iccma46720.2019.8988753-Figure1-1.png",
        "caption": "Fig. 1. North-East-Down ONED and Aircraft Body Center frame OABC co-ordinate systems [9]",
        "texts": [],
        "surrounding_texts": [
            "The quadroter co-ordinate system and configuration is represented as follows: Where the ONED and OABC represents the fixed and the body reference systems respectively. In the following section, the dynamic equations of the non linear model were derived by using Euler\u2019s and Newton\u2019s law in the literature. [9]"
        ]
    },
    {
        "image_filename": "designv11_14_0000454_j.jmapro.2015.12.003-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000454_j.jmapro.2015.12.003-Figure3-1.png",
        "caption": "Fig. 3. The fabricated process of a vertica",
        "texts": [],
        "surrounding_texts": [
            "154 C. Yanpu / Journal of Manufacturing\nNotation\nTd droplet temperature Ts substrate temperature Ps spraying pressure Wt puls width Hd deposition distance Oc oxygen content solidification angle d nozzle diameter D droplet diameter H formed height of single deposited droplet Hs slice thickness vp platform velocity W scanning step Wd formed maximum width Rb spreading radius of droplet Rc spherical cap radius\nt b c w m s t s e e p d fi o t w f g h a s fi f\nLn layer number\nogether to fabricate three-dimensional complex components layer y layer [10]. Because the metal droplets fusing together to form omponents is a complicated fluid and thermal behavior process, hich includes impacting, remelting, cooling and solidifying of etal droplets [11], it is very hard to precisely control the forming ize of parts. Many researchers have investigated the influence of he technical parameters on the forming shape of component after olidification [12] and bonding between metal droplets [13]. Howver, selecting an appropriate layer thickness is very important for ffectively controlling forming precision under the fixed process arameters. During fabricating metal components by sequentially epositing micro-droplets layer by layer, the 3D model of part rstly shoud be read into the slice software and sliced into a series f parallel layers with a certain slice thickness [14]. If the selecion of slice thickness is not appropriate, the fabricated part surface ould appear the\u201cstep phenomenon\u201dand the dimension error of ormed part would be increasing [15]. Like others AM technoloies, the surface curvature variation [16] of part and the formed eight of single deposited layer are two key factors in selecting the ppropriate slice thickness [17]. In the process of MDDM, when the urface curvature of fabricated part and the process parameters are xed, the selection of slice thickness should be matched with the ormed height of single deposited layer. If the selecting value of slice\nProcesses 21 (2016) 153\u2013159\nthickness (Hs) is too samll, the deposition layer number of part model after slicing and the forming dimension of part in deposition direction are increased. Contrary, if the selecting value of slice thickness (Hs) is too large, the deposition layer number of part model after slicing and the forming dimension of part in deposition direction are reduced. So, it is very important to study the influences of slice thichness on the forming dimension, and propose a convenient and effective method to ascertain an appropriate slice thickness (Hs) for obtaining good and desired component dimensions.\nIn this paper, a convenient and effective method was proposed to resolve the selecting problem of slice thickness. A series of deposition experiments were carried out to validate the selection method. The work provides an effective method for optimizing slice thickness and improving the dimensional accuracy of the formed components by droplet-based freeform fabrication.\n1. Process principle and experimental setup\nFig. 1 shows schematic diagram of process principle and experimental setup of metal micro-droplet deposition manufacture. The experimental setup was developed, which mainly included a dropon-demand generator, a droplet deposition system, a temperature measurement and control system, a process monitor system, and an inert environment control system. The ejection of droplets and the motion of deposition substrate are controled by experimental system according to the AM procedure file, and the metal part is fabricated by sequentially depositing metal droplets layer by layer.\nIn metal droplets deposition experiments, a pneumatic droplet generator was used to produce molten droplets on demand, which consisted of a droplet controller, a solenoid valve, a graphite crucible, a heating furnace and a nitrogen gas resource. The crucible was designed to be heated up to 1000 \u25e6C to melt metal alloy material by induction heating apparatus. A high graphite nozzle was embedded into the bottom of crucible and nitrogen gas was delivered to the top of crucible via the solenoid valve. Pulsed pressures were generated by opening and closing the solenoid valve alternatively, which was driven by the droplet controller. Uniform micro metal droplets were ejected out through the nozzle by the pulsed pressure. The temperature measurement and control system was used to measure and control the temperatures of metal droplets and deposition substrate, which consisted of high sensitivity thermocouples and temperature control meter (Yatai N800). The process monitor system consisted of a high speed CCD camera (Uniq 610), a microscope (Optem Zoom 65) and an image acquisition card. The camera was triggered to capture images of droplets ejection\nrinciple and experimental setup.",
            "l colu\na u g\n2\ni 3 i t t t n i\ni f s s F t H i d d\no l d f s t p t s\nf\nb radius Rc. Fig. 4(c) gives two kinds of cross-sections of solidified droplets with different solidification angles According to the conservation of mass, the volume of a spherical droplet with droplet\nnd deposition. The inert environment control system, which was sed to protect molten metal from oxidization, was made up of love-box and gas circulating device.\n. Theoretical analysis and calculation\nIn process of fabricating metal parts by sequentially depositng molten metal droplets on substrate layer by layer. First, the D model of part must be read into the slice software and sliced nto a series of parallel layers with a certain slice thickness. So, he selection of slice thickness is very important, which influence he forming dimension and deposited layer numbers. Fig. 2 shows he slicing results of a thin-walled tube under different slice thickesses. It can be seen that, the deposited layer number is obvious ncreasing with the decreasing of slice thickness (Hs). In metal micro-droplet deposition manufacture, it is a difficult ssues to select an appropricate slice thickness(Hs) to ensure the ormed dimension of part is cosisitent with the designed dimenion, when the process parameters (such as:droplet diameter, olidification angle, deposition temperatue and so on) are fixed. ig. 3 shows the fabricated process of a vertical colum by sequenially depositing droplets. It can be seen that, the forming height\nof the vertical colum is gradually heightening with the increasng of deposited layer number, and the formed thickness of single eposited layer is relevant to the solidified shape of deposited roplet on the substrate.\nBecause the selection of slice thickness is affected by the shape f solidified droplet on the substrate, it is very important to anayze the relationship of shape parameters of solidified droplet. The roplet was a nearly sphere with diameter D after being ejected rom the nozzle. Then the droplet impacted and solidified on the ubstrate, which was considered as a spherical cap The deposiion topography of single droplet on substrate is influenced by rocess parameters such as substrate temperature, deposition disance, oxygen content of environment and surface characteristic of ubstrate.\nFig. 4(a) shows the shape of droplets on substrate under diferent process parameters. Fig. 4(b) shows the cross-section of a\nmn by sequentially depositing droplets.\nsolidification droplet, which can be represented by shape parameters such as solidification angle, formed maximum width Wd, formed height H, spreading radius of droplet R and spherical cap",
            "t solid\ni s n h i i s D\nIn process of fabricating metal parts by sequentially depositng molten metal droplets on substrate layer by layer, the selected lice thickness shoud be consistent with the deposited layer thickess. So, the slice thickness Hs is approximately equal to the formed eight H of single deposited droplet(Hs = H). From Eqs. (4) and (5), t can be seen that the formed height H can be controlled by solidfication angles and droplet diameter D. Thus, it is known that the election of slice thickness Hs is only related to the droplet diameter\nand solidification angle . The appropriate Hs can be calculated\nification angle (a) = 132.5\u25e6 , (b) = 116.5\u25e6 and (c) = 92.5\u25e6 (d) = 55\u25e6 .\nby Eqs. (4) and (5) when knowing the D and of droplets. In order to validate the relationship between the formed height H and the solidification angle , the droplet diameter D is fixed, Fig. 5 shows the deposited results of a vertical column under different droplet solidification angles . The experimental results show that the formed height H of single deposited layer is obviously decreasing along with the diminishing of the .\n3. Experimental results and discussion\n3.1. Deposition experiment of thin-walled tube\nA series of experiments were conducted using Sn60\u2013Pb40 alloy under different Hs in order to validate the calculating model of slice thickness. Table 1 lists the process parameters of deposition experiment. Fig. 6 shows the Surface topography of metal droplets and deposited thin-walled tube. The 3D model of deposition sample (thin-walled tube) is shown in Fig. 7(a), the height of sample is designed to be 9.5 mm in deposition direction (in z direction). The copperplate was used as deposition substrate in experiments. The slice thickness Hs of sample 1, 2, 3, 4 and 5 was selected as 263 m, 306 m, 320 m, 365 m and 400 m, respectively. In the deposition experiments, the appropriate slice thickness Hz is 320 m according to Eqs. (4) and (5) when droplet diameter D is 400 mm, and solidification angle is 121.51. Fig. 7(b)\u2013(f) shows the deposited results of thin-walled tubes under different slice thickness Hs.\nIt can be seen that the formed dimension of the deposited samples is very sensitive to the slice thickness in deposition direction, which is gradually diminishing with the increasing of the slice thickness in z direction. Fig. 8 shows the deposited layer numbers Ln of samples under different slice thicknesses Hs. Fig. 9 shows the measuring results of the deposited samples under different slice thicknesses Hs. The formed dimensions of thin-walled tubes are measured by universal toolmaker\u2019s microscope, with the measurement method of average for many times.\nWhen Hs is selected as 263 m and 306 m, the 3D model of thin-walled tube was sliced into 36 and 31 layers, respectively. The sample 1 and 2 were fabricated under different deposited layer numbers, the formed dimension of samples are 11.50 mm and 9.82 mm, which are obvious bigger than the designed size (9.5 mm), and the relative error of samples are 21.05% and 3.36%,"
        ]
    },
    {
        "image_filename": "designv11_14_0001412_j.mechmachtheory.2019.103730-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001412_j.mechmachtheory.2019.103730-Figure7-1.png",
        "caption": "Fig. 7. A Schatz-based 7R mechanism with a tangential intersection in the configuration space.",
        "texts": [
            " Note that dim (\u03c02 ( \u0304K 4( j) q 0 )) = 2 = 2 dim (\u03c01 ( \u0304K 4( j) q 0 )) , j = \u03b1, \u03b2, therefore \u03ba2 = 4 and dim (\u03c02 ( \u0304K 4( j) q 0 )) will no longer decrease and the vector x 2 will remain unchanged through the solution sets of any order. Observe that q\u0308 7R (q 0 , x 1 ) = x 7 = 0 in branch \u03b1 and q\u0308 7R (q 0 , x 1 ) = x 7 = (32 / 5) (x 1 ) 2 = 0 in branch \u03b2 , which 2 2 1 proves that the seventh R joint is active in \u03b2 . Fig. 6 shows the 7R mechanism in two configurations each belonging to each motion branch. 6.2. Case 2: a Schatz-based 7R mechanism Fig. 7 a shows a 7R mechanism in which joints with axes S 1 , S 2 , S 1 , S 3 , S 2 and S 4 constitute a Schatz 6R linkage [63,64] . As shown in Fig. 7 b, in the configuration q 0 \u2208 V, axes S 1 , . . . , S 4 lie on plane , while axes S 1 and S 2 lie on and are perpendicular to . A seventh R joint is inserted between joints with axes S 3 and S 2 . At q 0 , the axis of the seventh joint, S , is parallel to S and also lies on . 7R 1 The screw coordinates with respect to the coordinate system with origin at O shown in Fig. 7 a are the following: S 1 ( q 0 ) := ( 0 , \u2212 1 , 0 ; 0 , 0 , 0 ) , S 2 ( q 0 ) := ( \u22121 , 0 , 0 ; 0 , 0 , 0 ) , S 1 ( q 0 ) := ( 0 , 0 , 1 ; 2 , 0 , 0 ) , S 3 ( q 0 ) := ( 1 2 , \u221a 3 2 , 0 ; 0 , 0 , 1 ) , S 7R ( q 0 ) := ( 0 , 0 , 1 ; 1 2 , \u2212 3 \u221a 3 2 , 0 ) , S 2 ( q 0 ) := ( 0 , 0 , 1 ; 0 , \u2212 2 \u221a 3 , 0 ) , S 4 ( q 0 ) := ( 0 , \u2212 1 , 0 ; 0 , 0 , \u2212 2 \u221a 3 ) , Define x i := (x 1 i , . . . , x 7 i ) \u2208 R 7 , where x 1 i := d i q 1 / d t i , x 2 i := d i q 2 / d t i , x 3 i := d i q 1 / d t i , x 4 i := d i q 3 / d t i , x 5 i := d i q 7R / d t i , x 6 i := d i q 2 / d t i and x 7 i := d i q 4 / d t i "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000808_cgncc.2016.7829020-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000808_cgncc.2016.7829020-Figure1-1.png",
        "caption": "Figure 1. Earth inertial and body-fixed frames of quadrotor.",
        "texts": [
            " Section will draw a conclusion. The quadrotor is equipped with four rotors in a cross configuration. There are two pairs of propellers which are placed on the end of the cross-like structure. To maintain the balance of the overall torque, one pair of rotors rotates in a clockwise direction while the remaining pair rotates in a counter-clockwise. There are two modes in accordance with the orientation of the motion which are \"x\" mode and \"+\" mode. The \"+\" mode used in the present paper and the quadrotor is shown in Fig. 1. And in this figure two reference 978-1-4673-8318-9/16/$31.00\u00a92016 IEEE frames should be defined to derivate the dynamics, namely, the earth inertial frame {E} and the body-fixed frame {B} which are used to describe the relative motions between the coordinate frames. For the convenience of discussion in this section, we would like to take the following assumptions into account. Assumption 1: The body is rigid and symmetrical. Assumption 2: Both the body-fixed frame origin and the center of mass of quadrotor are assumed to coincide"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001055_j.jfluidstructs.2019.06.005-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001055_j.jfluidstructs.2019.06.005-Figure1-1.png",
        "caption": "Fig. 1. Geometric parameters of the swimmer and the channel, representations of rotating magnetic field B, the gravity vector g and channel inlet flow with a parabolic profile and average velocity, vf . Forward (head direction, pusher-mode) and backward (tail direction, puller-mode) motion of the swimmer.",
        "texts": [
            " In addition to the viscous force, the external magnetic torque, the gravity force and the normal contact force on the swimmer are also considered in the CFD model. Simulations are carried out to study the effects of the Mason number, which is the ratio of viscous and magnetic torques, the diameter ratio of the swimmer and the channel, tail length and the channel flow on swimmer trajectories. In addition to improving the fundamental understanding, these results bear strong relevance to the development of control systems for microswimming robots. The geometric setup is shown in Fig. 1 where the swimmer with a left-handed helical tail and a cylindrical head with curved edges is placed inside a cylindrical channel of diameter Dch. The length of the swimmer\u2019s tail is denoted by L, wavelength by \u03bb, amplitude by Bsw and the diameter by dtail. The length of the cylindrical head is Lhead and the diameter is Dhead. The radius of curvature of the edges is rc. The channel length is set to a very low but acceptable value, which is almost twice as long as the length of the swimmer, and helps to reduce the computation time while the end effects on the swimmer remain negligible",
            " The swimmer geometry is representative of the swimmers used in our experiments (Acemoglu and Yesilyurt, 2015; Caldag et al., 2017), and it is also similar to many others used in the literature (Ghosh and Fischer, 2009; Tottori et al., 2012). Swimmers are identified with the letter \u2018\u2018L\u2019\u2019 followed by a number that represents the number of waves on the tail. Pusher and puller swimming modes, which are defined based on the position of the head with respect to the tail and the swimming direction, are also depicted in Fig. 1. For a left-handed helical tail, the pusher-mode swimming occurs when the swimmer rotates in the counter-clockwise direction and the puller-mode in the clockwise direction. A fully-developed Poiseuille flow with a parabolic profile and average velocity vf is specified at the inlet; vf = 0 for the base case. Gravity g acts in the negative y-direction. Fluid motion is governed by the steady Stokes equations at low Reynolds numbers as time-dependent effects such as the history and added mass forces are negligible as long as the rotation frequency, f, is not very high (Wang and Ardekani, 2012): \u2207 2u \u2212 \u2207p = 0, \u2207 \u00b7 u = 0 (1) Here, u and p are the nondimensional fluid velocity field and the pressure, respectively",
            " The viscous force is obtained from the integration of fluid stress on the swimmer: Fvj = \u222b S \u03c3ijnidS (4) where \u03c3ij are the elements of the stress tensor for i = 1, 2, 3 and j = 1, 2, 3, ni denotes the ith component of surface normal, S is the swimmer surface, and summation over repeated indices is implied. The viscous torque exerted by the fluid with respect to the center of mass of the swimmer is: \u03c4vj = \u222b S (xs \u2212 x) \u00d7 \u03c3ijnidS (5) where xs is the position of a point at the surface and x is the position of the center of mass of the swimmer. The magnetic field rotates around the x-axis (Fig. 1), which is also the centerline of the channel, and exerts a magnetic torque on the swimmer: \u03c4m = m \u00d7 B (6) where m is the magnetization vector of the swimmer with a magnitude of m0. The rotating field is achieved by out-of-phase sinusoidal fields and given by: B = B0 [ 0 cos (\u03c9t) sin (\u03c9t) ]\u2032 (7) where B0 is the amplitude of the magnetic field, \u03c9 = \u00b12\u03c0 f is the rotation rate and its sign implies the rotation direction of the left-handed helical tail that pushes the swimmer when \u03c9 > 0 and pulls it when \u03c9 < 0",
            " Thus, the velocity of a point on the swimmer surface is given by: us = U + \u03c9 \u00d7 (xs \u2212 x) (10) One end of the channel is the inlet where the fully developed Poiseuille flow with a parabolic velocity profile is specified, and the other end of the channel is defined as the outlet where the pressure is set to 0. Trajectories of the swimmer are obtained from the kinematic relations: dx dt = U (x, ei) (11) dei dt = \u03c9 (x, ei) \u00d7 ei (12) The ei for i = 1, 2, 3 represent the unit vectors of the local coordinate system placed at the center of mass of the swimmer, as shown in Fig. 1, and form the columns of the rotation matrix, which are used to calculate the Euler angles to fully define the swimmer orientation in the global channel coordinates. Since the inertial effects are negligible, acceleration of the swimmer is not considered. Linear and angular velocities are obtained from the solutions of the steady Stokes equation by the CFD model at each position and rotation, hence the velocities only depend on the position and the rotation of the swimmer represented by the unit vectors, ei"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002865_10667857.2021.1913318-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002865_10667857.2021.1913318-Figure5-1.png",
        "caption": "Figure 5. Principle diagram of friction and wear.",
        "texts": [
            " The hardness change of the coating along the depth direction of the cross-section was tested by the HX-1000 Vickers microhardness tester. During the measurement, the load was 200 g and the loading time was 15 s. The Brookfield UMT3 M-220 multifunctional friction and wear machine (ball speed 100 r/min, loading load 10 kg, circular wear scar diameter 8 mm, friction time 30 min) was used to test the tribological properties of the coating. The friction mode was ball disc motion friction, and the friction pair was WC ceramic ball. The principal diagram of friction is shown in Figure 5. VHX-5000 ultra-fine optical microscope (OM) was used to observe the wear profile and wear scar morphology of the coating after friction. The corrosion resistance of the coating was studied by electrochemical corrosion experiments. The coating was polished, polished and cleaned, and the part except the surface was covered with epoxy resin and tested with an Autolab electrochemical workstation. In the experiment, the counter electrode was a platinum plate, the reference electrode was saturated, and the sample after laser cladding was the working electrode"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000158_j.mechmachtheory.2019.103625-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000158_j.mechmachtheory.2019.103625-Figure1-1.png",
        "caption": "Fig. 1. Conception of 1-dof planar linkage, (a) a 3-RRR 3-dof planar linkage, (b) a ten-bar 1-dof planar linkage in which input links are constrained by two parallelograms.",
        "texts": [
            " The linkages are designed by constraining 3-RRR planar parallel mechanisms and constructed with 10 bars. For this new linkage, synthesis problem is formulated with an algorithm developed. Input\u2013output (IO) equations are derived. This paper extends the research work reported by Bai [24] with refined formulation and analysis. Synthesis examples are included to illustrate the synthesis and analysis of the linkage. The new synthesis approach takes advantage of planar parallel mechanisms. A planar multi-dof mechanism can visit infinitive many poses within its workspace. Fig. 1 a shows a 3-RRR planar linkage. This is a 3-dof mechanism, where three input links B i C i , i = 1 , 2 , 3 are mounted to the ground via revolute joints labelled as B i . The coupler, or the mobile platform, is a ternary link, which connects three limbs at three revolute joints A i . The middle joints of each limb are labelled as C i . In the mechanism shown in Fig. 1 a, three driving angles are noted by \u03b8 ij , where j is the pose number. The mechanism can be converted to 1-dof planar linkage, if we make all three inputs \u03b8 ij identical. This can be achieved by adopting two parallelogram linkages, namely, B 1 B 2 C 2 G and B 2 B 3 HC 2 , on the two kinematic loops (Loops 1 and 2) of the parallel mechanism, which leads to a linkage with 13 revolute joints, as shown in Fig. 1 b. In the new linkage, there are two more ternary links, B 1 C 1 G and B 3 C 3 H used for the construction of two parallelograms. The input angles are thus expressed as \u03b8i j = \u03b80 i + \u03b8 j , where \u03b80 i is the angle of the i th limb at a reference position. As such, there is only one rotation variable \u03b8 j for all three input links so the mechanism has only one dof. The same conclusion can be made by applying the Gr\u00fcbler formula. The mechanism in Fig. 1 b is now used for exact motion synthesis. As the coupler link moves, it visits m given poses of the coupler link, described by r j and \u03c6j , j = 0 , . . . , m . All displacements are given with respect to the reference pose given by r 0 = 0 and \u03c60 = 0 . At the reference pose, the rotation angle \u03b8 j = \u03b80 = 0 . It is noted that mechanism in this synthesis is obtained by constraining 3-RRR planar parallel linkage. By this way, the number of degree of freedom is reduced from 3 to 1. Moreover, the obtained 10-bar planar mechanism has a smaller design space than the ordinary 10-bar linkages [25] , due to the presence of constraining parallelograms",
            " Without loss of generality, the planar parallel mechanism that is used for synthesis is of arbitrary configuration, not necessarily symmetric. For dyad A 1 C 1 , by virtue of the link rigidity, the length remains constant during motion. This means (a 1 j \u2212 c 1 j ) T ( a 1 j \u2212 c 1 j ) = l 2 1 , j = 1 , . . . , m (1) where a 1 j and c 1 j are position vectors of joints A 1 and C 1 at the j th pose, l 1 is the link length of A 1 C 1 . The equation is applicable to other two dyads, A 2 C 2 and A 3 C 3 , of Fig. 1 , which lengths l 2 and l 3 being constant too. So Eq. (1) is rewritten as (a i j \u2212 c i j ) T ( a i j \u2212 c i j ) = l 2 i i = 1 , 2 , 3 (2) where, a i j = r j + Q j a 0 i (3a) c i j = b i + R j c \u2032 i (3b) c \u2032 i = c 0 i \u2212 b i (3c) where Q j = Q (\u03c6 j ) and R j = R (\u03b8 j ) are matrices of planar rotation, while a 0 i and c 0 i are position vectors of joints A i and C i at the reference pose. Moreover, the lengths of three distal links A i C i are determined from the reference positions as l i = | a 0 i \u2212 c 0 i | i = 1 , 2 , 3 (4) Eq",
            " , 9 , we finally obtained 18 equations containing design variables only. Assembling them together yields a system of 18 polynomial equations for 18 variables. The system admits exact solutions to generate linkages able to visit all 10 prescribed poses. The same conclusion can be reached if Eq. (10a) is used for synthesis, with extra variables \u03b8 j . Once all coordinates of joints in the reference pose are found from the synthesis equations, coordinates of points G and H for constructing two parallelograms are readily found. Taking the parallelograms of Fig. 1 as example, the coordinates of points G and H in vector form, namely, g and h , are g = b 1 + c \u2032 2 ; h = b 3 + c \u2032 2 ; (17) The ten-bar 1-dof linkage is different from the four-bar linkage and other linkages due to the constrained input angles. An input\u2013output (IO) equation is needed for position and branch analysis. The IO equation can be obtained using the forward kinematics of the 3-RRR parallel mechanism, which has been well documented in literature [26,27] . The reported methods are mainly based on the loop closure equation involving the input rotations and also the end-effector coordinate and orientation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001483_s42835-020-00365-1-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001483_s42835-020-00365-1-Figure5-1.png",
        "caption": "Fig. 5 SMA connector for coaxial probe feeding",
        "texts": [
            " To obtain the circular polarization in the antenna, the length and width of the patch is tuned to form a dimension of square patch as 45\u00a0m \u00d7 45\u00a0mm, the substrate and ground dimensions as 60\u00a0mm \u00d7 60\u00a0mm using the CST Simulation software. The truncation at the corners of the square patch is to obtain the circular polarization of the antenna. The structure shown in the Fig.\u00a04 is similar to simple square patch antenna with compact size. The antenna is fabricated on a 3.2\u00a0mm thickness (i.e.) h = 3.2\u00a0mm FR4 substrate is used. The Coaxial feed dimension is calculated to match 50\u00a0Ohm impedance. The Fig.\u00a05 shows the SMA connector to fabricate the antenna. The SMA connector is designed in the CST software using the impedance calculator for the coaxial connector and also determined the dimensions as, Inner feed outer diameter = 0.56\u00a0mm Teflon outer diameter = 1.9\u00a0mm Outer Shield diameter = 2.1\u00a0mm The discrete port is connected between the inner feed and outer shield to transmit the power into the antenna. The discrete port is selected by clicking Simulation \u2192 discrete port option. (11)RL(dB) = 20 log | | 1 3 In this design of GPS-GLONASS antenna, the researchers used the FR4 substrate which is a glass-reinforced epoxy laminate sheets used in Printed Circuit Boards (PCBs)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure8-1.png",
        "caption": "Fig. 8 Comparison of rake face",
        "texts": [
            " And the 3D model is generated as shown in Fig. 6. Different from design angles (or called shape angles), the working angles are dynamic in regard to the cutting velocity, which directly affects the cutting process. The working angles defined in the orthogonal plane can better reflect the cutting performance of the cutter. Taking the workpiece in Section 2.4 as an example, the cutter with a plane rake face is also designed as shown in Fig. 7, and the comparison with the curved rake face is shown in Fig. 8. For convenience, the cutter with curved rake face is named cutter A, and the cutter with plane rake face is named cutter B. It is obvious that the working rake angle of cutter A is just equal to \u03b3o since the rake face takes the working rake angle as the design variable directly. However, the working rake angle of cutter B is not distinct since its rake face is designed according to the shape rake angle \u03b3e. For comparison, the working rake angle of cutter B is calculated by use of the reference system as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003872_j.1460-2687.1999.00029.x-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003872_j.1460-2687.1999.00029.x-Figure1-1.png",
        "caption": "Figure 1 Theoretical case of a rigid sphere impacting on a rigid surface.",
        "texts": [
            " The \u00aerst objective of this paper was to describe quantitatively the dynamic behaviour of cricket balls impacting on simulated pitch constructions and to assess the variations in behaviour due to the use of different soils and grasses. The second objective was to assess how the results compared with other pitch measurements (e.g. impact hardness, as measured by Lush 1985) and soil characteristics (e.g. moisture content). The theoretical case of a rigid sphere impacting on a rigid surface was examined by Daish (1972). Figure 1 shows a rigid sphere travelling from the left with spin imposed upon it (vertical velocity, vvi, is taken as being positive into the ground and back spin, xi, is positive). After impact with a rigid horizontal surface the sphere moves to the right but remains in the same plane (vertical velocity, vvo, is taken as being positive away from the ground and top spin, xo, is positive). During impact, the left-to-right movement is opposed by a frictional force, F. There is also a reactive force, R, acting vertically upwards with the sphere having mass M and radius a"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002535_s00170-020-06330-6-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002535_s00170-020-06330-6-Figure2-1.png",
        "caption": "Fig. 2 Wide-beam laser and powder flow model",
        "texts": [
            " Based on the wide-beam laser energy distribution and powder concentration distribution, and the geometrical positional relationship between laser beam and powder flow, a wide-beam laser-powder interaction model in lateral powder feeding laser claddingwas constructed. This interactionmodel includes laser energy attenuation model and powder particle temperature distribution model. Furthermore, the effect of main parameters on the interaction model was discussed. Figure 1 shows the experimental equipment of the wide-beam laser cladding, including a 3 kW continuous wave high-power diode laser, a nozzle with square outlet (13 mm \u00d7 1 mm), an ABB robot, and a worktable. Figure 2 depicts the schematic of laser and powder beam during wide-beam laser cladding. As shown in Fig. 2, the powder nozzle is located on the front of the laser scanning direction. OLxLyLzL is the laser coordinate system, and OL is fixed at the intersection of the laser beam and power beam axis. And OLxLyLzL is chosen as the reference frame Oxyz. Opxpypzp is the powder flow coordinate system, which the origin is located in the powder outlet center. Before establishing the interaction model of wide-beam laser and powder flow, the energy distribution of the widebeam laser and the concentration distribution of powder flow must be analyzed first"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003092_s10999-021-09559-5-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003092_s10999-021-09559-5-Figure7-1.png",
        "caption": "Fig. 7 Finite element model of a rectangular membrane under shear loading with zoomed-in crease profile",
        "texts": [
            " Moreover, the crease profile considered here is circular and dependent on neutral angle relaxation. Difference observed here is not significant. Selection of the correct packaging method is an essential task considering membrane shape-size and stowage requirements. Despite the advantages and limitations of available techniques: wrapping, origami folding; Z-folding is adopted as a baseline here due to its simplicity. Material properties definition considered in the test case are introduced in a baseline configuration of single creased laminates. The geometry profile shown in Fig. 7 is prepared with the assumption of flat film, i.e. membrane angle is zero and a constant radius of curvature. However, aspect ratio, pre-stress value and other parameters are kept constant. Though it was expected to establish correct properties, adhesive layer properties are neglected for numerical simulation. Mesh should be dense enough with at least six nodes over each wrinkle wavelength. This may not be achieved all the time for practical computational timing. Hence, meshing is done with 30,794 thin shell elements (S4R5), including 2076 in a creased portion, maintaining sufficient density in all regions with more refinement near the crease to observe accurate results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002849_s13369-021-05659-8-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002849_s13369-021-05659-8-Figure1-1.png",
        "caption": "Fig. 1 View of commercial IM",
        "texts": [
            " In this section, the commercial 4\u00a0kW IM is analyzed using FEM and the simulation results are validated using experimental data. In the design process of an electric machine, the operating temperature is an important factor. The temperature rise of the different components of 4-pole induction motors (4-55\u00a0kW) has been presented in [21, 22]. A 2D model of the commercial IM was developed using the stated motor dimensions. The stator and rotor magnetic 1 3 material were chosen to be non-oriented steel (M47050A). The main specifications of the motor were determined from the values given in Table\u00a01. Figure\u00a01 illustrates some views of the commercial IM under study. The LSPMSM design procedure was presented in detail in previous work [21]. The design methodology is based on combining the rotors of an IM and an interior permanent magnet (IPM) motor. The stator and winding for both the IM and IPM motor are the same. The configuration of the permanent-magnet locations and the volume and rotor bar shape are calculated based on sizing equations to reach a balance between starting and synchronization performance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000745_0731684416678670-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000745_0731684416678670-Figure8-1.png",
        "caption": "Figure 8. Preparation of specimen with different scale of fiber micro-buckling: (a) sketch of experimental set-up and (b) specimen with micro-buckled fibers prepared by this set-up.",
        "texts": [
            " Moreover, longitudinal tensile properties of lamina are dramatically influenced by in-plane fiber micro-buckling, while the influences of in-plane fiber micro-buckling on transversal tensile properties can be ignored. Preparation of specimen with different scales of in-plane fiber micro-buckling To verify the results obtained in \u2018\u2018Macroscopic characterization of in-plane fiber micro-buckling\u2019\u2019 and \u2018\u2018Theoretical analysis of the influences of in-plane fiber micro-buckling on tensile properties\u2019\u2019 sections, specimen with different scales of micro-buckled fiber are prepared based on the formation mechanism of fiber micro-buckling in AFP process. The experimental set-up used in this article is shown in Figure 8(a). The mechanism of this experimental setup can be described as below: prepreg tow is placed on the surface of a pre-stretched elastomer. After a large enough adhesive stress is generated between prepreg tow and pre-stretched elastomer, tensile deformation of elastomer will disappear and the axial compressive deformation will be generated in prepreg tow when the tension acted on elastomer is released. Thus, fiber micro-buckling is formed and the scale of it depends on the tensile strain of elastomer. Specimen with fiber micro-buckling prepared by this experimental set-up are shown in Figure 8(b). The prepreg tow used in this article is EH104/D12 slit tows with 6.35mm in width and 0.15mm in Figure 6. Typical stress distributing graphs of different scaled in-plane fiber micro-buckling with same displacement: (a) \"trajectory\u00bc 0, (b) \"trajectory\u00bc 0.3%, (c) \"trajectory\u00bc 0.9%, and (d) \"trajectory\u00bc 1.2%. Table 1. Details of EH104/D12 slit tows. Ex (GPa) Ey (GPa) Gxy (GPa) nxy XT (MPa) XC (MPa) YT (MPa) YC (MPa) S (MPa) at University of Otago Library on November 22, 2016jrp.sagepub.comDownloaded from thickness, supplied by Hengshen Co"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002904_tmech.2021.3074800-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002904_tmech.2021.3074800-Figure1-1.png",
        "caption": "Fig. 1. (a) Overview of the assistive walker. (b) Side view of the assistive walker: (1) active wheels, (2) bevel gears, (3) DC servo motors, (4) motor drivers, (5) base frame, (6) casters, (7) battery, (8) Industrial Personal Computer, (9) linear actuator, (10) support bar, (11) links, (12) parallel support frame, (13) handles, (14) support plate, (15) hand guard.",
        "texts": [
            " The main contributions of this study are summarized as follows: 1) a simple and effective assistive walker to assist the user in walking and STS transfer; 2) an optimized ANFIS based intelligent STS intention recognition approach, which outperforms conventional methods; 3) a safe, stable and compliant control algorithm, which encourages the user to actively participate in STS transfers; and, 4) real-world experimental verifications of the proposed control algorithm with young healthy and elderly subjects. This section introduces the assistive walker and its sensory system for measuring the interaction forces and estimating the user\u2019s posture. We designed an assistive walker as shown in Fig. 1, that consists of the following three subsystems: \u2022 A mobile platform, which was realized by differential driving of two rear wheels. \u2022 An STS support subsystem, which used only one linear actuator (GJ20-05-120, Shanghai Guang Jian automation equipment Ltd, electrical parameters in Table I) to provide assistive force/torque during the STS transfer and to maintain the height of the interaction module. Compared Authorized licensed use limited to: Makerere University Library. Downloaded on May 17,2021 at 14:36:18 UTC from IEEE Xplore"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001350_0954407019890481-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001350_0954407019890481-Figure4-1.png",
        "caption": "Figure 4. Water jacket CFD model. CFD: computational fluid dynamics.",
        "texts": [
            " Figure 3 shows the CFD-FEA mapping results in the first cylinder under rated speed condition, including gas temperature and HTC computed by 3D combustion CFD. The other three cylinders share the same HTBC as the first cylinder. The cylinder head has much higher HTC than other components. It also has a very high gas temperature (averaged over the whole cycle). Its peak temperature is higher than the one of the exhaust port. Thus, it has the highest thermal load among the components of the engine. The water side of the FEA- CFD coupled interface refers to the water jacket surface of the cylinder head and block. Figure 4 shows the water jacket CFD model used in analysis. The cooling water jacket serves as the cooling boundary of the engine cylinder head. The total number of grid elements is about 5.2million. The water Figure 2. In-cylinder combustion CFD analysis model. CFD: computational fluid dynamics. jacket ensures that the cylinder head is at the appropriate temperature without failure. Meanwhile, thermal load of the cylinder head also depends on the homogeneity of temperature distribution. Higher temperature gradient will result in higher thermal load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001412_j.mechmachtheory.2019.103730-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001412_j.mechmachtheory.2019.103730-Figure8-1.png",
        "caption": "Fig. 8. Branches of motion of the Schatz-based 7R mechanism.",
        "texts": [
            " The solution set for the third order constraints is: K\u0304 3 q 0 = {( x 1 , x 2 , x 3 )\u2223\u2223x 1 \u2208 \u03c01 ( K\u0304 2 q 0 ) , x 2 \u2208 \u03c02 ( K\u0304 2 q 0 ) , x 3 = ( x 1 3 , \u221a 3 8 ( 4 x 1 3 \u2212 ( x 1 1 )3 ) + 9 8 x 1 1 x 5 2 , 3 \u221a 3 2 x 1 1 x 1 2 \u2212 1 4 x 5 3 , \u221a 3 4 ( 4 x 1 3 \u2212 ( x 1 1 )3 ) , x 5 3 , 3 \u221a 3 4 x 1 1 x 1 2 \u2212 3 4 x 5 3 , \u22123 \u221a 3 8 x 1 1 x 5 2 + 3 4 ( x 1 1 )3 + 1 2 x 1 3 ) x 1 1 , x 1 2 , x 5 2 , x 1 3 , x 5 3 \u2208 R } , The solution set for the forth order constraints is: K\u0304 4 q0 = K\u0304 4(\u03b1) q0 \u222a K\u0304 4(\u03b2) q0 where, K\u0304 4 ( \u03b1) q 0 = { ( x 1 , x 2 , x 3 , x 4 ) \u2223\u2223x 1 \u2208 \u03c01 ( K\u0304 2 q 0 ) , x 2 = ( x 1 2 , \u221a 3 2 x 1 2 , \u221a 3 2 ( x 1 1 )2 , \u221a 3 x 1 2 , 0 , \u221a 3 4 ( x 1 1 )2 , 1 2 x 1 2 ) , x 3 = ( x 1 3 , \u221a 3 8 ( 4 x 1 3 \u2212 ( x 1 1 )3 ) , 3 \u221a 3 2 x 1 1 x 1 2 \u2212 1 4 x 5 3 , \u221a 3 4 ( 4 x 1 3 \u2212 ( x 1 1 )3 ) , x 5 3 , 3 \u221a 3 4 x 1 1 x 1 2 \u2212 3 4 x 5 3 , 3 4 ( x 1 1 )3 + 1 2 x 1 3 ) , x 4 = ( x 1 4 , \u221a 3 4 ( 2 x 1 4 \u2212 3 ( x 1 1 )2 x 1 2 ) + 3 2 x 1 1 x 5 3 , \u221a 3 4 ( 7 ( x 1 1 )4 + 6 ( x 1 2 )2 \u2212 8 x 1 1 x 1 3 ) \u2212 1 4 x 5 4 , \u2212 \u221a 3 2 ( 3 ( x 1 1 )2 x 1 2 \u2212 2 x 1 4 ) , x 5 4 , \u221a 3 8 ( 7 ( x 1 1 )4 + 6 ( x 1 2 )2 + 8 x 1 1 x 1 3 ) \u2212 3 4 x 5 4 , \u2212 \u221a 3 2 x 1 1 x 5 3 + 9 2 ( x 1 1 )2 x 1 2 + 1 2 x 1 4 ) , with x 1 1 , x 1 2 , x 1 3 , x 5 3 , x 1 4 , x 5 4 \u2208 R } and K\u0304 4 ( \u03b2) q 0 = { ( x 1 , x 2 , x 3 , x 4 ) \u2223\u2223x 1 \u2208 \u03c01 ( K\u0304 2 q 0 ) , x 2 = ( x 1 2 , \u221a 3 2 x 1 2 , 0 , \u221a 3 x 1 2 , 2 \u221a 3 ( x 1 1 )2 , \u22125 \u221a 3 4 ( x 1 1 )2 , 1 2 x 1 2 ) , x 3 = ( x 1 3 , \u221a 3 8 ( 4 x 1 3 + 17 ( x 1 1 )3 ) , 3 \u221a 3 2 x 1 1 x 1 2 \u2212 1 4 x 5 3 , \u221a 3 4 ( 4 x 1 3 \u2212 ( x 1 1 )3 ) , x 5 3 , 3 \u221a 3 4 x 1 1 x 1 2 \u2212 3 4 x 5 3 , \u2212 3 2 ( x 1 1 )3 + 1 2 x 1 3 ) , x 4 = ( x 1 4 , \u221a 3 4 ( 15 ( x 1 1 )2 x 1 2 + 2 x 1 4 ) + 3 2 x 1 1 x 5 3 , \u221a 3 4 ( 16 ( x 1 1 )4 + 6 ( x 1 2 )2 + 8 x 1 1 x 1 3 ) \u2212 1 4 x 5 4 , \u2212 \u221a 3 2 ( 3 ( x 1 1 )2 x 1 2 \u2212 2 x 1 4 ) , x 5 4 , \u221a 3 8 ( 25 ( x 1 1 )4 + 6 ( x 1 2 )2 + 8 x 1 1 x 1 3 ) \u2212 3 4 x 5 4 , \u2212 \u221a 3 2 x 1 1 x 5 3 + 1 2 x 1 4 ) with x 1 1 , x 1 2 , x 1 3 , x 5 3 , x 1 4 , x 5 4 \u2208 R } It can be seen that \u03c01 ( \u0304K 4(\u03b1) q 0 ) = \u03c01 ( \u0304K 4(\u03b2) q 0 ) and \u03c02 ( \u0304K 4(\u03b1) q 0 ) = \u03c02 ( \u0304K 4(\u03b2) q 0 ) , therefore q 0 is a tangential intersection of V \u03b1 and V \u03b2 with n C = 1 . Since dim (\u03c02 ( \u0304K 4( j) q 0 )) = 2 = 2 dim (\u03c01 ( \u0304K 4( j) q 0 )) , j = \u03b1, \u03b2, therefore \u03ba2 = 4 . Observe that q\u0308 7R (q 0 , x 1 ) = x 5 2 = 0 in branch \u03b1 and q\u0308 7R (q 0 , x 1 ) = x 5 2 = 2 \u221a 3 (x 1 1 ) 2 = 0 in branch \u03b2 , which proves that the seventh R joint is active in \u03b2 . Fig. 8 shows the 7R mechanism in two configurations each belonging to each motion branch, where geometric constraints lead to metamorphic behavior that limits the motion of links. 7. Conclusions This paper presented the first examples of reconfigurable mechanisms whose configuration space contains tangent inter- sections of two branches of motion of the same dimension. It was pointed out that these singularities cannot be analyzed by computation of the kinematic tangent cone. In these examples, the dimension of both branches is one"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000122_j.colsurfa.2019.123928-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000122_j.colsurfa.2019.123928-Figure2-1.png",
        "caption": "Fig. 2. The strategy adopted for shape transformation.",
        "texts": [
            " Bending angle, height, and end-to-end curve distance were measured as depicted in Fig. 1 (b). All the analysis was done in triplicates. Designing shape-shifting products with homogeneous density distribution throughout the matrix require a complex and continuous spatial force gradient thereby making the process practically very complicated. Moreover, isotropic swelling does not result in a change of the shape [36]. On the other hand, a product with inhomogeneous density distribution swells anisotropically resulting in shape change as illustrated in Fig. 2. The density gradient in the xerogel discs was created by adopting the method of sessile drop drying of the wheat hydrogel. During drying there will be an accumulation of more particles at the top than bottom thereby making top denser than the bottom. This might be due to the formation of gelled skin or crust on the top surface when the polymer reaches a critical concentration during evaporation of polymer solution at a temperature below its glass transition [37]. The inhomogeneous density distribution induces anisotropic swelling resulting in bending"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000876_0954407018824943-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000876_0954407018824943-Figure3-1.png",
        "caption": "Figure 3. Automatic tensioning device and its schematic diagram. FEAD: front engine accessory drive.",
        "texts": [
            "11 Particularly, determining Wbelt-h-bending requires, for each pulley of the FEAD some integral calculations, over the belt cross section for each layer or rubber composing the belt as the bending strain depends on the pulley radius and the distance from the pitch line. The tensioner-hysteresis characterizes the power dissipated by friction during the tensioner-arm movement. The tensioner-arm/pulley is a part of the automatic tensioning device whose function is to adjust the belt slack span tension value over a wide range of operating conditions (Figure 3). Tensioner pivot is realized by a plain bearing which induces friction and consequently energy loss. Indeed, when the tensioner-arm moves, it causes a relative movement in the pivot assembly and thereby generates friction and thus damping (Figure 4). According to Michon et al.,15 the dissipated energy due to the tensioner hysteretic behaviour equals to the internal area of the hysteresis loop in Figure 4. This amount of energy lost can be calculated by equation (6) Wtens h =2TfcDu \u00f06\u00de where Wtens h is the dissipated energy in the tensioner pivot in (J); Du, the amplitude of the tensioner rotation angle in (rad); and Tfc, the tensioner frictional torque in (Nm)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000067_metroi4.2019.8792876-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000067_metroi4.2019.8792876-Figure1-1.png",
        "caption": "Fig. 1. The satellite antenna mechanism and the belt support studied.",
        "texts": [
            " Requirements of the measurement techniques are discussed in order to obtain measurements realiable for the validation of the simulation model and for the understanding of the physical phenomea involved in the process. The methods able to measure the effect of residual stress on a macro-scopic scale will be considered, in particular CMMs and other optical measurement techniques, based on laser displacement sensor. The need of post-processing will be also discussed with reference to the integration of measurements and model data. II. MATERIALS AND METHODS The test case refers to a part, which is the support of the antenna of a satellite, in close collaboration with Thales Alenia Space as seen in Fig.1. This part started from an initial conventional design and was re-formed several times using topological optimization tools aiming to reduce the total weight of the component and maximize, in parallel, its rigidity. This process includes several steps like stress-analysis (external loads) and topological optimization (re-design) of it with respect to these specific loads. Both the original and the optimized components may be seen in Fig.2. and Fig.3, respectively. The material of this part is the Inconel 625 nickel-iron alloy which is frequently used in additive manufacturing process (PBF)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002746_j.triboint.2021.106920-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002746_j.triboint.2021.106920-Figure15-1.png",
        "caption": "Fig. 15. Spiral grooved texture.",
        "texts": [
            " The average leakage was equal to the flowrate during the outstroke minus the flowrate during the instroke divided by two. Fig. 14 shows that the circular grooved texture had little impact on O-ring leakage under current operating conditions. In the calculation results of the various groove parameters, the leakage amount changed by only 2.63% at most. Fig. 16 shows the transient fluid film pressure and transient asperities contact pressure of the reciprocating rod seal with a spiral grooved texture rod (see Fig. 15). The results are like those for circular grooved texture rod to some extent. The difference is that the fluid film pressure was lower and the asperities contact pressure was higher, but both pressures fluctuated more. The reason for this phenomenon is that the spiral groove forms a pressure relief channel, which results in a sharp drop in fluid film pressure and the virtual height at equilibrium decreases (see Fig. 17). Therefore, the fluctuation of fluid film pressure and asperities contact pressure increases because a lower film thickness yields a pressure that is more sensitive to the change in film thickness"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000264_012031-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000264_012031-Figure1-1.png",
        "caption": "Figure 1. The position of specimen for Izod (a) and Charpy (b) impact test.",
        "texts": [
            " Generally, there are four types of impact velocity which are low velocity, high velocity, ballistic and hypervelocity [1-3]. The Charpy and Izod impact test were categorized in low velocity impact, which determine the impact strength and toughness of a material. An impact blow is delivered to a test specimen by means of a pendulum-type hammer and the impact value is determined from the energy required to break the specimen. Results of impact tests are expressed in term of impact energy (J/m) or impact strength (kJ/m2). Figure 1 illustrates the experimental set up for Izod and Charpy impact test. ICADME 2019 IOP Conf. Series: Materials Science and Engineering 670 (2019) 012031 IOP Publishing doi:10.1088/1757-899X/670/1/012031 It was shown that both Charpy and Izod impact test applied almost the same principle. Only several differences, such as the shape and position of the specimen may differ the type of impact test. Table 1 summarize the differences between Charpy and Izod impact test. In industrial scale, Charpy impact can be described as an economical quality control method to determine the impact toughness and notch sensitivity of materials"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003870_(sici)1097-0363(19990815)30:7<845::aid-fld867>3.0.co;2-o-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003870_(sici)1097-0363(19990815)30:7<845::aid-fld867>3.0.co;2-o-Figure1-1.png",
        "caption": "Figure 1. Adjustable hydrodynamic bearing concept.",
        "texts": [
            " Performance characteristics predicted by the computer model have been demonstrated in practice. Herein is described the comprehensive mathematical model of the novel bearing and an automated computational process devised to produce solutions for a variety of simulated operating conditions. It was designed and developed to be applicable to all forms of the bearing and was used extensively to study, and optimise, the design of a particular type of the inverse arrangement comprising a rotor supported by a stationary journal. Figure 1 shows this Copyright \u00a9 1999 John Wiley & Sons, Ltd. Int. J. Numer. Meth. Fluids 30: 845\u2013864 (1999) in outline form with displacements greatly exaggerated for clarity. There are four adjustable cantilevered segments, G1, G2, G3 and G4, each supported by adjuster pins labelled \u2018A\u2019. The principle of the bearing was that the hydrodynamic conditions could be changed by independently controlled adjustments of the adjustable segment shapes and positions. This was effected by position change inputs to the adjuster pins, \u2018A\u2019"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure17-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure17-1.png",
        "caption": "Fig. 17. Principle of Scoop-Controlled Coupling",
        "texts": [
            " In view of the low heat dissipation from the stationary casing and the low head created by the pumping disk, however, a cooler of large area is required, and this type of coupling has not been generally adopted. Scoop-Controlled Coupling. By far the most successful self-contained coupling as yet produced is the \u201cscoop-controlled coupling,\u201d having a radially movable scoop tube within a rotating chamber enlarged sufficiently to contain the whole of the liquid required for the operation of the coupling. .This is shown in Fig 16, Plate 3, and Fig. 17. The working circuit is of the interrupted core-ring type, and leak-off nozzles are arranged at the periphery in the usual manner. The outer casing is largc enough to receive centrifugally the full contents of the working circuit, which forms a rotating annulus of the depth indicated by the dotted line. When stationary the liquid is contained in the lower half of the casing and there is no need for an external reservoir or liquid-tight glands. The scoop tube housing is mounted on a plain bracket, the scoop itself being pivoted so that it can swing through an angle of about 70 deg"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001594_j.cma.2020.112996-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001594_j.cma.2020.112996-Figure1-1.png",
        "caption": "Fig. 1. Initial configuration of the rod. (a) Reference frame (red), (b) material frame (black). (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " From the above geometry description Eq. (1), the initial configuration of the rod in the abstract space is C = { (r, d1, d2, d3) : L = (0, L) \u2192 R3 \u00d7 S2 \u00d7 S2 \u00d7 S2, di \u00b7 d j = \u03b4i j } (2) where \u03b4i j is the Kronecker delta, and the tangent d1 (S) is d1(S) = d2(S) \u00d7 d3(S) = r\u2032(S)/ r\u2032(S)  (3) with \u201c\u00d7\u201d indicating the cross product of vectors, (\u00b7)\u2032 differentiation with respect to S, and \u2225\u00b7\u2225 the Euclidean norm. For Kirchhoff rods, the reference frame { dr i (S) } and rotation angle \u03c6 are used to describe the material frame {di (S)}, as in Fig. 1. The smallest rotation (SR) reference frame [11] (or the natural reference frame in [14]) is defined in terms of the reference frame {dr i (0)} at S = 0 and the tangent d1 (S) as dr \u03b1 (S) = \u039b [d1(0), d1(S)] dr \u03b1 (0) (4) The rotation operator \u039b (n0, n) is interpreted as the rotation from the initial unit direction vector n0 to the current unit direction vector n that yields the minimum rotation angle as \u039b (n0, n) = (n0 \u00b7 n) 1 + n\u03020 \u00d7 n + 1 1 + n0 \u00b7 n (n0 \u00d7 n) \u2297 (n0 \u00d7 n) (5) where 1 is the identity tensor and the \u201c\u02c6\u201d operator maps a vector to a skew-symmetric matrix such that a\u0302b = a \u00d7 b, \u2200a, b \u220b R3",
            " For the Kirchhoff rod of interest, the vanishing shear strain condition leads to the relation \u03d1 = \u03c6d1 + d1 \u00d7 u\u2032 (12) where the angle of twist \u03c6 is related to the tangent d1 as \u03c6 = \u03d1 \u00b7 d1 (13) Consequently, the reduced space of infinitesimal displacement for the Kirchhoff rod can be uniquely defined by the axis displacement u and angle of twist \u03c6: V = { (u, \u03c6) \u2208 R3 \u00d7 R|u = 0 on \u0393u; \u03d1 = 0 on \u0393\u03d1 } (14) Remark 1. For general nonlinear cases, the Kirchhoff constraint is automatically met by the reference frame and rotation angle representation of the material frame, see [12] for more details. Through this representation, the rod\u2019s configuration can be defined by four degrees of freedom, i.e., the Cartesian components of the position r (S) and the rotation angle \u03d5(S) from the reference frame to the material frame about the tangent d1 (S) (see Fig. 1). The SR reference frame is chosen herein, due to its compatibility with the classical linear theory for the three-dimensional Kirchhoff rod, namely, the infinitesimal increment of the rotation angle \u03d5 with reference to the SR frame is consistent with the angle of twist \u03c6 in the classical linear theory, see Section 4.1 in [14] for proof. Since the angle of twist Eq. (13) is connected to the axis geometry, the conventional angle-of-twist interpolation fails to capture the rigid-body modes associated with infinitesimal rotations, as will be discussed in the isogeometric formulation to follow"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002307_itnec48623.2020.9084796-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002307_itnec48623.2020.9084796-Figure2-1.png",
        "caption": "Fig. 2. Clamp is designed to hold probe .",
        "texts": [
            " On the one hand, UR5 complies with point 5.10.5 of the standard EN ISO 10218-1:2006 which means UR5 is credible in Human-Robot interaction, and on the other hand, the UR5 with 6 degrees of freedom(DOF)can reach an arbitrary position with at least one pose within a working radius of up to 33.5 ins (850mm). The standard Denavit\u2013Hartenberg(DH)[13] model of UR5 and a calibrated parameters is provided by manufacturer. A handling gripper is designed to affixing the ultrasound probe on the distal end of the arm (Fig. 2). As a programmable, economic haptic device, Touch has attracted many researchers\u2019 attention in the field of robot control[14]. Based on T Sansanayuth et al.\u2019s effort[15], we summarize standard DH parameters as shown in Table \u2160, where 1 2 0.13335l l m  . In master-slave motion control, a control strategy is proposed as shown in Fig. 3. Mater\u2019s position pm needs to be filtered and s represents a time derivation of pm, and then Touch\u2019s final linear velocity vm is calculated. kv means scaling factor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002046_icem49940.2020.9270988-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002046_icem49940.2020.9270988-Figure1-1.png",
        "caption": "Fig. 1 IPMSM 48s8p topology from Toyota Prius 2004",
        "texts": [],
        "surrounding_texts": [
            "\u03a6Abstract -- This paper studies the effect of uneven pole magnetization on acoustic noise and vibrations radiated by Permanent Magnet Synchronous Machines (PMSM) under Maxwell force excitations in open-circuit condition and on the full speed range.\nFirst, the effect of magnetic asymmetries is analyzed using analytic equations based on permeance / magnetomotive force. Then, electromagnetic, structural and acoustic calculations are performed using MANATEE software on a 48s8p Interior PMSM used in Electric Vehicle (EV) / Hybrid Electric Vehicle (HEV) automotive applications. To speed up calculations under asymmetrical conditions due to magnetization variations, Electromagnetic Vibration Synthesis algorithm is used. Simulation results show that uneven pole magnetization significantly affects airborne noise and vibrations, inducing low frequency noise and new resonances with structural modes which are not predicted by the Greatest Common Divider (GCD) rule on pole/slot combination.\nIndex Terms\u2014electrical machines, powertrain, vibration, acoustic, e-NVH, uneven magnetization, manufacturing tolerances, numerical simulation\nI. INTRODUCTION The design of electric powertrains for EV/HEV applications includes optimizing Noise, Vibrations and Harshness (NVH) performances, especially those due to the electromagnetic forces generated by the electric motor (emotor) and referred as e-NVH. Electric motors are designed to have high power density, which also means that their structure is subject to high magnetic stress. Such high power density can be achieved by choosing permanent magnets as rotor excitation which can lead to major issues in terms of eNVH. Besides, the intrinsic variable speed property of powertrains imply additional e-NVH risks due to switching noise and the fact that resonance probability increases with speed range. To lower e-motor dimensions, rotating speeds are indeed pushed to their upper limits in recent Electric Drive Units (EDU).\nFurthermore, several e-NVH risks needs to be addressed at design stage. A detailed transfer path analysis of the powertrain can help to identify main e-NVH risks [1]: \u2022 airborne noise due to vibrations of the outer structure\nwhich propagate to the air. Airborne noise then arises from the direct excitation of the outer structure, generally by radial forces applying at the interface between stator and airgap. It generally occurs at \u201chigh frequency\u201d (1-7 kHz).\n\u2022 structure borne noise from various origins, the main ones being excitations (such as Unbalance Magnetic Pull\n\u03a6E. Devillers, P. Gning and J. Le Besnerais are with EOMYS ENGINEERING, 121 rue de Chanzy, Lille-Hellemmes, France (website: www.eomys.com, e-mail: contact@eomys.com).\nAmong e-motor topologies used in EV/HEV powertrains, the Interior Permanent Magnet Synchronous Machine with Zs = 48 stator slots and 2p = 8 poles (hereafter noted IPMSM 48s8p), as initially used in Toyota Prius 2004 powertrain, has become one of the main IPMSM design and especially because this particular topology demonstrates only a few eNVH risks among those detailed earlier. In fact, the 3-phases integral distributed windings associated to the high pole number enables to lower the risk of exciting radial modes by considering the well-known GCD rule, with GCD(Zs,2p) = 8 sufficiently high so that only pulsating forces may excite the structure and generate noise, especially within some specific speed ranges where those pulsating forces resonate with the breathing mode [2][3].\nHowever, powertrains produced in large series generally present other e-NVH risks due to manufacturing tolerances and faults, including rotor eccentricity [4], stator bore ovality [5], magnet displacement in slots, and uneven magnet magnetization [6]. This article proposes to further study the impact of uneven pole magnetization on the e-NVH performances of the IPMSM 48s8p topology at open-circuit (no-load) condition. Uneven magnetization distribution along rotor bore radius can be due to dispersion of the remanent flux provided by magnet supplier, uneven rotor temperature, or magnet displacement in slots during operation.\nIn this paper, the remanent flux of the different IPMSM poles are varied to study the impact of rotor uneven magnetization on stator magnetic vibrations and resulting\nEffect of uneven magnetization on magnetic noise and vibrations in PMSM\n\u2013 application to EV HEV electric motor NVH E. DEVILLERS, P. GNING, J. LE BESNERAIS\nAuthorized licensed use limited to: University of Exeter. Downloaded on May 30,2021 at 07:13:47 UTC from IEEE Xplore. Restrictions apply.",
            "airborne noise. MANATEE software [7] specialized in emotor electromagnetic, structural and vibroacoustic simulations is used to perform all e-NVH simulations and post-processings. Variable-speed acoustic results (spectrograms, order tracking, deflection shapes) are discussed and analyzed.\n1) Theoretical force analysis Electromagnetic forces considered in this study are Maxwell forces applying on stator structure and concentrated at iron/air interface. Maxwell force distribution, or Maxwell stress, is decomposed in elementary travelling harmonic waves of electrical frequency f and wavenumber r noted {f,r}. The frequency f is the frequency of force, vibration and acoustic noise. The wavenumber is the spatial frequency (i.e. number of maximums or minimums) of the stress wave and its sign indicates the rotation direction of the force wave.\nStress waves with wavenumber r = 0 are pulsating waves, whose frequency is proportional to Least Common Multiple (LCM) between slot and pole numbers, so LCM(Zs,2p)fs/p = 12fs (noted H48 in terms of mechanical order). Stress waves with wavenumber different of 0 are rotating waves. The lowest wavenumber is given by GCD(Zs,2p) = 8. Therefore, all airgap stress wavenumbers are either 0 or proportional to 8. All pulsating stress waves are surrounded by two rotating sidebands of frequency \u00b12fs (\u00b1H8) and wavenumber 8. This analysis is confirmed using MANATEE Campbell diagram feature based on slot/pole combination (cf. Fig. 4).\nUneven magnetization removes the periodicity given by the number of pole pairs: all magnetic poles are different. Therefore, all mechanical orders and wavenumbers can be present in Maxwell stress harmonic content.\n2) Flux density computation\nThe radial and tangential airgap flux densities in opencircuit condition are computed using MANATEE fast hybrid model coupling Subdomain Model (SDM) and Finite Element\nAuthorized licensed use limited to: University of Exeter. Downloaded on May 30,2021 at 07:13:47 UTC from IEEE Xplore. Restrictions apply."
        ]
    },
    {
        "image_filename": "designv11_14_0003883_37.569709-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003883_37.569709-Figure3-1.png",
        "caption": "Fig. 3. Our prototype manipulator cell: roller wheels (left) are arranged orthogonally to provide a force in any direction.",
        "texts": [
            " For example, if one cell breaks down, neighboring cells can detect the fault and work around the broken cell until it can be replaced. The application of the Virtual Vehicle is discussed in more depth in a previous paper by the authors [9]. System Architecture Mechanical Configuration We have built a prototype system consisting of a small array of cells capable of transporting objects about the size of a breadbox. The simple task of a cell is to create a directable force. Each cell consists of a pair of orthogonally oriented \u201croller wheels\u201d (see Fig. 3) which are capable of producing a tangent force and which allow free motion parallel to their axes. Each wheel is driven through a gear reduction by a small DC motor. Each of these cells is connected to a large breadboard style base (see Fig. 4) to create a regular array of manipulators. Current resources restrict us to build only 20 cells, which limits our capabilities in two dimensions. The mounting system allows us to reconfigure the setup into a single row, giving us the flexibility to study the mechanical and other physical aspects of a small two-dimensional grid and gain a clear grasp on communication and theoretical issues with a long one-dimensional row"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure12-1.png",
        "caption": "Fig. 12. Virtual prototype model of 3-R(RRR)R+R HAM.",
        "texts": [
            "1017/S0263574720001290 Downloaded from https://www.cambridge.org/core. University of Toledo, on 03 Jun 2021 at 19:25:40, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. desired trajectories, with an accuracy of 10\u22129. This confirms the accuracy of the inverse kinematic solutions. The 3D model of 3-R(RRR)R+R HAM is drawn by SolidWorks software, which is imported into Adams dynamics simulation software and added with constraints and motion joints. The virtual prototype model is shown in Fig. 12. The length of members Pi Ji , Ji Gi ,Gi Ki and Ki Qi is L = 0.21m, and the radius of the fixed platform O and the moving platform C is R = 0.15m. Given the material of each component in ADAMS software, the masses and moments of inertias of 3-R(RRR)R+R HAM are measured as shown in Table II. According to the pitching and azimuth trajectory and rotation characteristics of the polarization mechanism shown in Table I, ADAMS software is used to carry out dynamic simulation of the HAM and the obtained pitching and azimuth motion states are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001393_s40435-019-00605-x-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001393_s40435-019-00605-x-Figure3-1.png",
        "caption": "Fig. 3 Experimental autopilot system",
        "texts": [
            "Given the input spectrum\u03a6uu (\u03c9), output spectrum\u03a6yy (\u03c9), and cross spectrum\u03a6uy (\u03c9), the empirical frequency responseG (\u03c9) andmagnitude-squared coherence \u03a5 2 uy at a given frequency \u03c9 were calculated, respectively, as G (\u03c9) = \u03a6uy (\u03c9) \u03a6uu (\u03c9) (9) \u03a5 2 uy = \u2223\u2223\u03a6uy (\u03c9) \u2223\u22232 |\u03a6uu (\u03c9)| \u2223\u2223\u03a6yy (\u03c9) \u2223\u2223 (10) As a rule of thumb, the process linear operating range is defined over a frequency band for which \u03a5 2 uy \u2265 0.6 and the coherence function is not oscillating [31]. In this work, the experimental autopilot system is the one integratedwithAR.Drone 2.0 quadcopter platform, as shown in Fig. 3a. For this autopilot system, the input and output signals are accessible via Parrot Software Development Kit (SDK). The selected quadcopter system is powered by a threecell lithium-polymer GiFi battery providing a flight time of approximately 20min. The drone hardware includes a navigation board, two video cameras, a main board, and four geared brushless motors with each integrated with an electronic controller card. The navigation board contains the drone sensors, which are a 3-axis gyroscope, a 3-axis accelerometer, a 3-axis magnetometer, a pressure sensor, and an altitude ultrasound sensor. The video cameras are a VGA bottom camera and an HD frontal camera. The main board is responsible for processing the sensor data and video streams to achieve the desired motion. To communicate with the experimental autopilot system, a Windows laptop with a wireless link running at approximately 16Hzwas used as aGroundControl Station (GCS), as illustrated in Fig. 3b. The GCS employs two software components, an open-source C++ software module, same like the one used in [22], and MATLAB Simulink. The C++ program was used to decode the navigation data sent by the drone and to encode and send the navigation commands to the drone. The MATLAB Simulink was used for visualizing and recording the navigation data as well as for generating the navigation commands. For convenience, the C++ software module is modified such that the quadcopter motion is described using the standard aerospace convention (where the positive z-axis of a coordinate frame is pointed downward)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002352_j.mechmachtheory.2020.103945-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002352_j.mechmachtheory.2020.103945-Figure3-1.png",
        "caption": "Fig. 3. Lower modules: (a) \u201c \u02dc R l1  \u0303 R l2 \u201d, (b) \u201c \u02dc R l1 \u22a5 \u2194 R l2 \u201d, (c) \u201c \u02dc R l1 \u22a5 \u2194 R l2 \u22a5 \u0303 R l3 \u201d, (d) \u201c \u02dc R l1  \u0303 R l2 \u22a5 \u2194 R l3 \u201d and (e) \u201c \u02dc R l1  \u0303 R l2 / \u0303 R l3 \u201d.",
        "texts": [
            " By considering plane-symmetry and enumerating different joint layouts, topological structures of each kind of module can be synthesized, and the related results are presented as follows, in each of which the layouts of only R pairs in the left of the SP and in the SP. 2.2.1. Coupling chain modules There are three \u201c[3R]\u201d modules denoted as \u201c \u02dc R c1 / \u0303 R c2 \u201d, \u201c \u02dc R c1  \u0303 R c2 \u201d and \u201c \u2194 R c1 \u22a5 \u0303 R c2 \u201d, shown in Fig. 2 (a\u2013c). Besides, there are only one \u201c[2R]\u201d module denoted as \u201c \u02dc R c1 \u201d and only one \u201c[1R]\u201d module denoted as \u201c \u02dc R c1 \u201d, shown in Fig. 2 (d) and (e), respectively. 2.2.2. Lower modules There are two kinds of \u201c{4R}\u201d modules, \u201c \u02dc R l1  \u0303 R l2 \u201d and \u201c \u02dc R l1 \u22a5 \u2194 R l2 \u201d, shown in Fig. 3 (a) and (b), and there are three kinds of \u201c{6R}\u201d modules, \u201c \u02dc R l1 \u22a5 \u2194 R l2 \u22a5 \u0303 R l3 \u201d, \u201c \u02dc R l1  \u0303 R l2 \u22a5 \u2194 R l3 \u201d and \u201c \u02dc R l1  \u0303 R l2 / \u0303 R l3 \u201d, shown in Fig. 3 (c\u2013e), respectively. Please cite this article as: W.-a. Cao, Z. Jing and H. Ding, A general method for kinematics analysis of two-layer and twoloop deployable linkages with coupling chains, Mechanism and Machine Theory, https://doi.org/10.1016/j.mechmachtheory. 2020.103945 4 W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx Ru1 MP1 MP2 Ru2 Ru4 Ru3 Ru1 MP1 MP2 Ru2 Ru4 Ru3 Ru1 MP1 MP2 Ru2 Ru4 Ru3 Ru2 Ru1 Ru3 Ru4 MP1 MP2 (a) (b) (c) (d) W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx 5 Further, some units should be excluded due to their structural defects, such as local DOFs and undesired angle ranges[35]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000719_mma.4210-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000719_mma.4210-Figure2-1.png",
        "caption": "Figure 2. Trajectory of the end effector r. / expressed as a Hermite-type function and trajectory of the non-active end of the link.",
        "texts": [
            " The total length of the robotic arm denoted by r D dOP (distance between the manipulator base location O and its end-effector P/ varies because of the rotation of the rigid guide of Department of Design and Engineering, Faculty of Science and Technology, Bournemouth University, Poole, UK * Correspondence to: Mihai Dupac, Department of Design and Engineering, Faculty of Science and Technology, Bournemouth University, Poole, UK.. \u2020 E-mail: mdupac@bournemouth.ac.uk Copyright \u00a9 2016 John Wiley & Sons, Ltd. Math. Meth. Appl. Sci. 2016 the extensible manipulator. The end effector of the robotic arm OP should reach the points pi .xi ,yi/iD1,2,:::,n. The path followed by the non-active end S of the link SP controls the position of the active end of the robotic arm defined by the points pi .xi ,yi/iD1,2,:::,n shown in Figure 2. 2.2. Piecewise trajectory generation 2.2.1. Cartesian interpolating curves. In order to achieve some desired properties of the manipulator trajectory, piecewise polynomial interpolating curves may be considered. Such interpolating curves [2, 4, 13\u201317] guarantee slope continuity, and/or minimal data storage, and/or local control and smoothness (no abrupt changes in displacement and velocity). For the general case, Hermite interpolation is given based on derivative values at data points y .xi/ D yi , Py ",
            " A more complex representation generally having continuous first and second derivatives, which guarantee a smooth trajectory of the manipulator endeffector, is given by quantic Hermite interpolation in Equation (1) with mi 0 D yi , mi 1 D Pyi , mi 2 D Ryi , mi 3 D RyiC1 3Ryi 2hi 4PyiC1C6Pyi 10 yi h2 i , mi 4 D 3Ryi 2RyiC1 2h2 i C 7PyiC1C8Pyi 15 yi h3 i , mi 5 D RyiC1 Ryi 2h3 i PyiC1CPyi 6 yi h4 i . 2.2.2. Polar interpolating curves. To interpolate the periodic data frk , kgkD0,N, which define the rotating extending link with rk>08kand rN D r0, one can consider the strictly monotonic angles 0 < 1 < 2 < < N 1 < N D 0 C 2 defined on the real interval [ 0, 0 C 2 ]. In each interval f i , iC1giD0,N 1 defined by the two consecutive control points pi , piC1, a piecewise cubic interpolant (Figure 2) can be expressed as a Hermite-type function [3, 18] by r . / D 3X kD0 mi k . i/ k (2) where hi D iC1 i , yi D riC1 ri hi , mi 0 D ri , mi 1 D Pri , mi 2 D 1 hi \u0152 .2Pri C PriC1/C 3 ri , mi 3 D 1 h2 i \u0152Pri C PriC1 2 ri , ri D r . i/, riC1 D r . iC1/, and where the derivatives at the endpoints pi , piC1are Pr. i/ D dr. i/ d D Pri and Py . iC1/ D dr. iC1/ d D PriC1, respectively. One can parameterize the Cartesian coordinates using the polar coordinate q .r, /by q D r cos r sin D r cos sin D 3X kD0 mi k "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000514_j.protcy.2016.03.069-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000514_j.protcy.2016.03.069-Figure1-1.png",
        "caption": "Fig. 1. (a) Schematic diagram of elliptic bore bearing with coordinate system, (b) Representation of film by springs and dampers",
        "texts": [
            " It is worth noting here that ellipticity in the bearing bore arises due to the operational issues (fluctuation/variation in operating parameters and circularity imperfections in the shafts) and manufacturing constraints. Thus, the objective of this paper is to present numerical study for exploring the influence of ellipticity of the bore (with its orientation) and non-Newtonian rheology (thinning) of the lubricant on the performance and stability of bearings. 2. Mathematical model The non-circular bore bearing (having particular orientation with respect to the load line) with the coordinate system is shown in Fig. 1 along with representation of lubricating film by springs and dampers. 2.1. Film thickness relation Normalized film thickness is expressed as [2]: 2 21 cos cos 1 cos cos sinh G G X Y (1) where cosX , and sinY . For elliptic bore, the radial clearance Cr is expressed as (Rmin \u2013 R). Putting G = 0, eq. (1) reduces to the film thickness relation applicable for the circular bore. 2.2. Rheological relation The performance parameters of the fluid film journal bearings operating for wide range of loads and speeds are normally improved by adding the polymeric additives of high molecular weight in the lubricating oil"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001588_0954406220912786-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001588_0954406220912786-Figure3-1.png",
        "caption": "Figure 3. (a) FE model of irregular structure with a cube depicting the naming of faces, (b) FE model showing elements at junction and (c) FE model with increased strut thickness to represent as-built structures. FE: finite element.",
        "texts": [
            "49\u201351 Each test was conducted for five cycles, because stabilisation was achieved within five cycles during the initial testing of the samples. Finite element analysis and analytical modelling To study the effect of geometric deviation on the Young\u2019s modulus of the structure, a 5mm cube was used as the RVE of the structures for finite element analyses. The CAD file structures were meshed with HYPERMESH (HyperMesh V12, Altair Engineering Inc) using 10 node tetrahedron solid elements (SOLID 187), which were analysed using Ansys. The mesh model of the regular structure is shown in Figure 3. A mesh convergence analysis was carried out, and a mesh size of 0.06\u20130.08mm was used to obtain an average of 4,30,000\u20136,20,000 elements, depending on the strut thickness. The boundary conditions are described in Table 2. For regular structures, side 1 was loaded and side 3 was constrained in the vertical direction. The irregular and random structures were subjected to periodic boundary conditions, as listed in Table 2, where the nodes on side 4 and side 6 had coupled degrees of freedom. The processing and post processing of each analysis were performed using ANSYS (V16, Ansys Inc)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003837_s0165-0270(99)00020-5-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003837_s0165-0270(99)00020-5-Figure3-1.png",
        "caption": "Fig. 3. Scale drawing of the cutter. All home-made components are aluminum except the stainless plate next to the blade and thin copper plate holding the tiny plastic hammer. The distance between the blade and the small plastic hammer is precisely controlled by a micropositioner. The hammer-plate holder is omitted in the bottom view for simplicity. An asterisk indicates the point in the blade on which carbon fiber is positioned and the hammer hits.",
        "texts": [
            " An external motorizer controller box sets the speed and direction of motion of the upper aluminum plate. A full circuit diagram is available upon request. In outline, an appropriate region of the newly formed electrode is manipulated to touch the cutting blade (Schick, Injector Plus Platinum razor blade, WarnerLambert, Morris Plains, NJ) and a high-speed impact generates a clean uniform cut. The cutter comprises a blade holder and a miniature solenoid (Ledex\u00ae TDS, Lucas Control Systems, Vandalia, OH) serving as a hammer (Fig. 3). Both components are fixed to an L-shaped aluminum frame. The internal core of the solenoid advances rapidly when 12 V DC is applied through a microswitch. Because the core itself may rotate and quiver inside the coil, we placed the final hammer on a separate thin copper plate (0.2 mm thick) fixed to the same micropositioner as the solenoid to ensure reproducibility. It hits the carbon fiber and insulating plastic against the blade. The tiny hammer is simply the last 1.5 mm of a pipettor tip attached to the plate using a cyano-acrylate glue. To protect the hammer from being damaged by the sharp razor blade during impact, the inside and end were filled with silicone elastomer (Sylgard184, Dow Corning, Midland, MI) (Fig. 3, enlarged drawing in a circle). The position of the hammer relative to the razor blade is adjusted by a micropositioner so that the solenoid strike will cut only the carbon fiber but not the hammer itself. The position of the blade can be adjusted by shifting the blade adapter, held by screws in a U-shaped groove milled on a bottom plate (Fig. 3). The whole cutter is magnetically fixed on the stage of an inverted microscope (Fig. 4) and viewed through a 3.2\u00d7 objective lens. The X\u2013Y stage manipulator is used to bring the end of the electrodes into view. A long carbon fiber was cut several times to optimize the distance between the hammer and blade. The quality of cutting can be inspected at a magnification of 400\u00d7 . 2.4. Scanning electron micrographs Ends of several electrodes were cut and positioned on a conducting double-sided tape and metal stage",
            " (2) The duration of heating is set by a Fig. 4. Schematic drawing to show positioning of cutter and carbon-fiber electrode on an inverted microscope. A carbon-fiber electrode with long uncut fiber is held by an adapter (used as Adapter A in Fig. 1) that can swing. The movement adjusts the height of carbon fiber along the Z-axis. The X\u2013Y stage manipulator adjusts the position of the electrode in the plane parallel to the microscope stage. The cutter is fixed to the microscope stage by a magnetic tape (hatched objects in Fig. 3) attracted to a ferrous metal plate that is glued on the stage. Since the cutting edge of the blade has an angle of about 11\u00b0, the cut end of the carbon fiber can be pushed against the edge when the mini-hammer advances. To compensate for the angle, the cutter is rotated about an angle (u) of \u22126\u00b0 from the vertical line. The plastic hammer behind the screw is not well visible at the top view and not shown here. timer in the control unit. (3) Intermittent air puffs were used to form a blunt shank"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000245_j.jmatprotec.2019.116515-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000245_j.jmatprotec.2019.116515-Figure4-1.png",
        "caption": "Fig. 4. Finite element model of SSFC spinning: a) Geometry model; b) Mesh model of the blank.",
        "texts": [
            " In order to accurately obtain the thickness distribution of the hemispherical spun part, the three-dimensional contours of the spun parts were scanned by a portable 3D metrology arm FARO. Similarly, a three-coordinate measuring machine HEXAGON was used to obtain the section contour of the spun part to evaluate the shape deviation between the spun part and mandrel. 3. Simulation modelling and verification In this study, two finite element models of the SSFC and DSFC spinning with a blank diameter of 200 mm has been established respectively to simulate deformation behavior of the hemispherical parts. Fig. 4 shows the SSFC spinning simulation model based on ABAQUS/ Explicit software. In the model, the mandrel, tailstock and rollers were assumed to be analytical rigid bodies, and the blank was deformable material, whose mechanical parameters are shown in Table 2. In order to simulate the spinning deformation behavior of the material more truly, 3D 8-node reduced integration linear hexahedral element SC8R was mainly used to mesh the blank, as shown in Fig. 4b, and 3D 6-node reduced integration linear wedge elements SC6R was used as transition elements in the radial direction to control the aspect ratio of the elements. The size of element was controlled around one tenth of the working roller radius, and five thickness integration points are used. In the model, the working roller path is controlled by the coordinate transformation relative to the reference point, which is consistent with the actual spinning processing. The friction condition was assumed as followings: the friction coefficients between the blank and the mandrel was 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001665_s40436-020-00303-4-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001665_s40436-020-00303-4-Figure5-1.png",
        "caption": "Fig. 5 Marker\u2019s frame",
        "texts": [
            " The algorithmic steps are presented as Algorithm 1. Fig. 3 Point cloud segmentation H. Zhang et al. By applying Algorithm 1, the point clouds can be easily separated into three clusters, each of which corresponds to a single sphere (see Fig. 4). In order to track the position and orientation information of the moving marker, a coordinate system is set up according to the distribution of colored spheres. The position of the coordinate is placed at the center of the middle sphere; the pose of the coordinate is illustrated in Fig. 5. In this manner, the pose of the coordinate can be identified with respect to the origins of the three spheres. The point cloud of each sphere has been acquired; hence, the origin of each point cloud can be identified using data-fitting algorithms. Fitting a model to noisy data (regression analysis) is frequently employed in computer vision for a wide range of objectives, such as reverse engineering. As a traditional technique, the method of least squares has become the most popular algorithm applied in regression analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003727_1.2831176-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003727_1.2831176-Figure5-1.png",
        "caption": "Fig. 5 Schematic of spindle, cutter axis tilt and cutter offset geometry",
        "texts": [
            " Cutter center offset error refers to the situation where the cutter axis is displaced parallel to the axis of rotation by some amount. Such type of errors are typically found in setscrew type of tool holders. In general, both spindle and cutter axis tilt may occur in any arbitrary direction measured in a plane perpendicular to the ideal spindle axis. In this paper, the spindle tilt is assumed to have a magnitude and direction whereas the cutter axis is sup posed to be tilted only in the plane containing the cutter axis and the feed direction. Figure 5 illustrates the various coordinate systems and the process geometry variables that are influenced by the presence of all three of the aforementioned effects. The spin dle axis tilt magni tude and direction are represented by y,s and /3\u201e, respectively; the cutter axis tilt magni tude and cutter center offset are given by y,c and e,c. The posit ion of the j'th cutter tooth in the X'Y'Z' coordinate system, assuming that all three effects are present, can be found from the following equation, (Xl(t)) Yl(t) Zl(t) 1 > = COS P,s sin P,s 0 L 0 - s i n p,s 0 cos p,, 0 0 1 0 0 cos yts 0 0 sin y,s 0 1 0 0 - s i n 7\u201e 0 cos y,s 0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001752_j.mechmachtheory.2020.103992-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001752_j.mechmachtheory.2020.103992-Figure1-1.png",
        "caption": "Fig. 1. Coordinate systems in the meshing process.",
        "texts": [
            " Further, the curvature relationships between the tooth surfaces of PCSBGs are deduced based on tensor analysis and surface theory, which represent a more clear and intuitive process of obtaining relative normal curvature equations. Finally, logarithmic spiral bevel gears (LSBGs) are used as a case to represent the simple application of a kinematic geometry model and analytical equations of PCSBGs presented in this article. 2. Kinematic model of spiral bevel gear with continuous pure-rolling contact 2.1. Applied coordinate systems in meshing process A basic coordinate system is designed as shown in Fig. 1 . The Cartesian coordinate systems S p and S 0 are fixed in absolute space; the Cartesian coordinate systems S 1 and S 2 are fixed with the pinion and the gears, respectively, which means that their base vectors rotate with the revolution of the corresponding gears. The relative rotation angle between S 0 and S 1 is denoted by \u03c8 and the relative rotation angle between S p and S 2 by \u03c6. (1) and (2) are the tooth surfaces of the pinion and the gear, respectively. Similarly, (1) and (2) are the contact paths of the pinion and the gear, respectively, and M is the instantaneous contact point",
            ", v (12) = 0 = v (12) [ 1 ] = \u02d9 \u03c8 M 12 (\u03c8) d M 21 (\u03c8) d \u03c8 r (1) [ 1 ] (t) , (3) where M ij refers to the coordinate transformation matrix from S j to S i , and \u02d9 \u03c8 refers to the time derivative of \u03c8 , which is the rotation velocity of the pinion. Without loss of generality, let \u02d9 \u03c8 = 1 . A detailed representation of the coordinate transformation matrixes are provided in the appendix. Substituting Eqs. (2) into (3) and assuming \u02d9 \u03c8 = 0, the continuous pure-rolling condition becomes { t = \u03c8 p = i 21 sin \u03be \u22121+ i p cos \u03be , (4) 21 where \u03be refers to the shaft angle between z 0 to z p , as shown in Fig. 1 . Further, synthesizing Eqs. (3) and (4) results in the following equations: M 12 (t) d M 21 (t) d t r (1) [ 1 ] (t) = v (12) [ 1 ] \u02d9 t = 0 \u02d9 t = 0 d M 21 (t) d t r (1) [ 1 ] (t) = M 21 (t ) M 12 (t ) d M 21 (t) d t r (1) [ 1 ] (t ) = M 21 (t) v (12) [ 1 ] \u02d9 t = M 21 (t) 0 \u02d9 t = 0 . (5) From Eq. (4) , the position vector of the contact path of the gear, (2) , can also be represented in the homogeneous form as r (2) [ 2 ] (t) = M 21 (\u03c8) r (1) [ 1 ] (t) = M 21 (t ) r (1) [ 1 ] (t ) = \u239b \u239c \u239c \u239d f (t) sin ( i 21 t ) (p cos \u03be \u2212 sin \u03be ) f (t) cos ( i 21 t ) (p cos \u03be \u2212 sin \u03be ) f (t)( cos \u03be + p sin \u03be ) 1 \u239e \u239f \u239f \u23a0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000728_icelmach.2016.7732503-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000728_icelmach.2016.7732503-Figure1-1.png",
        "caption": "Fig. 1. Typical rotor structures of SynRM.",
        "texts": [
            " The operation of the machine relies on the natural trend of the low reluctance structures (d-axis) to align themselves with the magnetic field [10]. Torque production capability depends on the difference between the d-axis and q-axis inductances( Ld \u2013 Lq ) which is known as saliency. The operating power factor in a SynRM, in a given operating point, depends on the ratio of the d-axis to the q-axis inductances ( Ld/Lq ) which is known as saliency ratio [11]. The most commons types of SynRM are the salient pole rotor, the axially laminated rotor and the transverse laminated rotor (see Fig. 1) [12]. The most employed topology is the axially laminated rotor with air barriers, since it offers good saliency ratio (i.e. 5 to 10) with an acceptable manufacturing cost [5], [13]. The number of flux barriers and their thickness are the parameters that decide the saliency ratio and the performance [8]. 1) Theoretical model with constant parameters: A synchronous reluctance machine can be described by a set of equations in the rotating d-q reference frame attached to the rotor: vd = Rsid + d\u03bbd dt \u2013 \u03c9\u03bbq \u03bbd = Ldid vq = Rsiq + d\u03bbq dt + \u03c9\u03bbd \u03bbq = Lqiq (1) Where vd, vq are d-axis and q-axis voltages; id, iq are d-axis and q-axis currents; \u03bbd, \u03bbq are d-axis and q-axis components of stator flux linkages; Ld, Lq are d-axis and q-axis inductances; \u03c9 is the electrical rotor speed and Rs is the stator resistance"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002256_access.2020.2973341-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002256_access.2020.2973341-Figure1-1.png",
        "caption": "FIGURE 1. Mechanism for enhancement steering angle and active locomotion: (a) Configuration of the proposed mechanism, (b) Steering angle by a ball joint angle, and (c) principle of generation of the improved steering angle by the joint angle \u03b8J and the wire angle \u03b8W .",
        "texts": [
            " The rotating magnetic field corresponded to a more stable steering ability than the dc magnetic field did because the robot and wire were aligned in the direction of the rotation axis. Through various experiments, we have improved the steering angle and verified the active movement using the proposed mechanism at a low magnetic field strength. II. MECHANISM AND MANIPULATION OF THE PROPOSED SYSTEM A. BALL JOINT-BASED END-EFFECTRO FOR ENHANCEMENT OF STEERING We proposed a novel active guidewire mechanism that employs a ball joint mechanism with a spiral-type magnetic microrobot to improve the steering angle of active guidewire as well as active locomotion. Figure 1 shows the configuration of the proposed mechanism and the conditions for steering via the ball joint mechanism. The proposed active guidewire consists of a spiral-type magnetic microrobot, ball joint, and guidewire, as shown in Fig. 1 (a). The microrobot is synchronized via a rotating magnetic field. In particular, 31104 VOLUME 8, 2020 the ball joint connects the robot and the guidewire while the robot rotates to tow and steer the wire. In general, the steering angle for magnetically actuated guidewires is the result of the deformation of a guidewire, and active locomotion is not employed. Therefore, the generated steering angles are small, with values of less than 90\u25e6. However, because the developed active guidewire utilizes the ball joint mechanism, two steering angles corresponding to the deformation and ball joint angles are generated. In this manner, the proposed mechanism can provide steering angles of up to 150\u25e6. In addition, the spiral-type magnetic microrobot provides two functions: the realization of active locomotion and drilling in blood vessel to remove clots. The fabricated ball joint can provide a joint angle of up to 45\u25e6 along with no wire deformation, as shown in Fig. 1 (b). Therefore, when the desired angle \u03b8D is 45\u25e6 or less, the steering angle \u03b8S is the joint angle \u03b8J . If the desired steering angle is more than 45\u25e6, the steering angle of the proposed guidewire is the sum of the joint angle and wire angle (\u03b8W ), that is the total steering angle is \u03b8S = \u03b8J + \u03b8W , as shown in Fig. 1 (c). The proposed ball joint mechanism not only provides a wide steering angle but also provides precise control and tilting of the position of the end effector in the rotating magnetic field. The tilt angle is equal to the ball-joint angle. B. MAGNETIC MANIPULATION TO REALIZE STEERING AND ACTIVE LOCOMOTION Because of the presence of the ball joint in the end-effector of the proposed mechanism, the developed guidewire provides a wider range of steering angle than can be generated using general guidewire"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003687_a:1008389419585-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003687_a:1008389419585-Figure3-1.png",
        "caption": "Figure 3. Gear test set-up (1-motor; 2-couplings; 3-spindle; 4-test gears; 5,6-infrared rays receivers; 7-spindle; 8-couplings; 9-reducer; 10-couplings; 11-loader).",
        "texts": [
            " It is, however, difficult to solve the continuoustime equation (26) analytically; therefore, a numerical solution procedure is developed. The time discrete equations for Equation (26) are, respectively, Xa(ti+1) = 8(ti+1, ti)Xa(ti)+ Bd(ti)u(ti)+ Ld(ti)w(ti), (33) where 8(ti+1, ti) = exp[Fa(t)(ti+1 \u2212 ti )], (34) Bd(ti) = ti+1\u222b ti 8(ti+1, ti)B(\u03c4) d\u03c4, (35) Ld(ti) = ti+1\u222b ti 8(ti+1, ti)G(\u03c4) d\u03c4. (36) The time interval 1t (= ti+1 \u2212 ti) is important during the calculation, and can be decided by \u2016Fa(t)\u20161t \u2264 \u03b4, (37) where \u03b4 is a constant, depending on the calculation precision, and \u2016 \u00b7 \u2016 presents the norm of matrix. Figure 3 shows the structures of the experimental system, which consists of two spindle boxes, load system, variable frequency control system of speeds, motor, hydraulic system, etc. The spindles, with average diameter of \u03c6 100 mm, the rotational precision of \u03c6 0.0015 mm error and supported by static pressure bearings, have much larger stiffness than that of the teeth of gears. Thus, the tested gear system can be modeled as a system with one degree of freedom. The noise was measured under normal gear working condition by the sound pressure, and took the average of three measured values"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000212_mees.2019.8896501-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000212_mees.2019.8896501-Figure6-1.png",
        "caption": "Fig. 6. Calculated points (a) and the values of the total magnetic force modulus \u2013 the vector sum of the projections (b).",
        "texts": [
            " The difference in the method for calculating the dependences of magnetic forces on a rotor displacement for a radial AMB is as follows. The full nominal gap 2 r (on both sides of a collar) is evenly divided into (2n+1) levels in both the vertical and horizontal directions. Each (n+1)th level coincides with the central position of the rotor between the stator poles. Further, the rotor is shifted so that the center of mass of its cross section coincides with one of geometrically permissible levels horizontally and vertically. They are indicated by dots in Fig. 6. The total of such calculations is about (2n+1)2. A static electromagnetic analysis is performed and the magnetic forces in the x, y directions and the total vector are determined at this mutual position (Fig. 6b). The results of analyzes at different rotor positions for a radial AMB with an outer stator diameter of 0.1 m and an internal pole diameter of 0.06 m with a nominal radial clearance of r=1 mm, taking into account the control law, are shown in Fig. 6b and 7\u0430. Fig. 7b shows the three-dimensional stiffness characteristic. It is obtained by approximating the magnetic force values at calculated points by complete polynomials of two variables x and y, followed by a differentiation. The calculation results for the radial AMB (Fig. 6, 7), as well as for the axial AMB (Fig. 2, 3), indicate that a negative feedback in the CS (see Fig. 4) and the selected control law parameters ensure a centering of the rotor in the gap. IV. CONCLUSION AND DISCUSSION The paper proposes numerical finite element methods for determining the force and stiffness characteristics of AMBs for almost any configuration and design. Their use allows to perform numerical experiments for a preliminary assessment and a search for rational or optimal bearing parameters without performing time-consuming natural experiments"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000763_s10443-016-9565-5-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000763_s10443-016-9565-5-Figure4-1.png",
        "caption": "Fig. 4 Meso-scale voxel-based RVC of warp-reinforced 2.5D woven composites: a yarns, b matrix",
        "texts": [
            " (1) Binder warp is divided into two parts, as shown in Fig. 2b. One is curve segment which keeps intimate contact with weft, while the other is line segment. (2) The section of weft yarn is convex lens, while the section of warp and blinder warp are rectangular, as shown in Fig. 2c. (3) Both warp and weft are straighten. The dimension of RVC can be calculated as A \u00bc 2 10 Pweft ;B \u00bc 2 10 Pwarp \u00bc 2 10 Pblinder ;C \u00bc Twarp \u00fe 2b \u00f01\u00de Where Twarp(mm)is the thickness of single warp. b(mm) is the length of minor semi-axis of weft. According to Fig. 4b, the geometrical relationship can be expressed as lc \u00bc r \u00fe Twarp 2 \u03b81; a \u00bc r sin\u03b81; b \u00bc r\u2212r cos\u03b81; Sweft \u00bc 4 r2 \u03b81\u22122 r\u2212b\u00f0 \u00de a; ld cos\u03b82 \u00bc A 2 \u2212a\u2212 r \u00fe Twarp sin\u03b81; ld sin\u03b82 \u00bc 4b\u00fe Twarp \u00f02\u00de Where a(mm) is the length of major semi-axis of weft. The value of ab is measured from the actual observation, here is equal to 4. r is the drawing radius of convex lens. \u03b81 is the angle of curve segment of weft. \u03b82 is the dip angle of linear segment of weft. lc(mm) and ld(mm) are the length of curve segment and linear segment, respectively",
            " Meso-scale voxel-based method, which divides warp-reinforced 2.5D woven composite RVC into a lot of regular 3D grid of voxels, is proposed to investigate the effective mechanical properties. In this study, RVC voxel mesh is generated with TexGen software. The advantages of this methodology for modeling 2.5D woven composites are (1) quick and automatic generation of voxel meshes; (2) assigning automatically the material properties; (3) providing the mesh periodicity at the boundaries. A voxel mesh schematic view of yarns and matrix is shown in Fig. 4. Undoubtedly, the fiber volume fraction has a first order effect on the mechanical properties of composites. Moreover, it is crucial for the definition of the material orientation in the elements, especially for the blinder warp. Figure 5 illustrates the material orientation of the blinder warp, Noted that the geometrical characteristics of RVC is sufficiently described. Furthermore, RVC is not isolated from its adjacent RVCs in the composites, and the boundary effects from the adjacent RVCs should be taken into account"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000684_cjme.2016.0617.074-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000684_cjme.2016.0617.074-Figure2-1.png",
        "caption": "Fig. 2. Structure of the flexibly mounted ring",
        "texts": [
            " Considering the need of development of DGS, this paper first research the influence of gas film thickness disturbance on sealing performance, what\u2019s more, the effects of some key factors on gas film thickness disturbance are systematically investigated at the operational conditions of high-speed and high-pressure, and it would provide a theoretical guide for understanding DGS\u2019s working rules and advancing DGS\u2019s design method. 2 Models 2.1 Physical models Fig. 1 shows a schematic cross section of a S-DGS with a flexibly mounted stator, and this model refers to Ref. [8] which published in 2002, it considers axial vibration of shaft and installation deviation of rotor. when the shaft spins, the installation deviation of rotor will provide an angular excitation motion to stator. The spiral groove geometry is shown in Fig. 2(a), when the rotor revolves at a high speed, there is a large hydrodynamic pressure will be generated by the spiral grooves, which will help to separate the two seal faces and maintain a steady gas film thickness. But in most of the practical seals, the gas film thickness has a disturbance because of the rotor\u2019s axial runout and misalignment. Fig. 3(a) shows the relative position between stator and rotor, Fig. 3(b) shows the model of seal face kinematics, and in the seal dynamic tracking property analysis, the gas film, which have a certain stiffness and damping properties, is usually treated as a spring and damper system, and the stator can be thought of as being supported by it[3]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003891_0042-6989(95)00201-4-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003891_0042-6989(95)00201-4-Figure1-1.png",
        "caption": "FIGURE 1. Schematic drawing of the main features for the calibration procedure. An eye with two search-coils (C1, frontal coil; C2, lateral coil) is shown within the cube of the generating coils; ~p being the angle between both coils. The voltages induced in the search-coils are Ul and U2. The calibration cube containing three exactly orthogonal calibration coils (induced voltages Ux, Uy, Uz) and a short circuit section for measuring offset voltages (U0) is positioned as close as possible in front of the eye. Note that different components are not to scale. Magnetic field directions indicate that fields do not have to be orthogonal to the coil frame. For further details see text.",
        "texts": [
            " Two search-coils (frontal and lateral), which were intentionally not exactly perpendicular, and a laser point were attached to the eye. The eye could be moved to all positions, including quarternary, with a gimbal system. The calibration procedure requires two essential steps: (1) measurement of the magnetic field by a calibration cube; and (2) calibration with spontaneous eye movements; to further improve the calibration; fixation of several targets can be used optionally (3). 2.2.1. Step 1: Measurement of the magnetic fields. The magnetic fields near the eye are measured with a special calibration cube (Fig. 1). This cube contains three exactly orthogonal calibration coils to measure the three spatial components of each magnetic field and a short circuit section to measure the offset voltages. To consider changes of the fields by the introduction of the subject or experimental devices, the cube measurement is done with the monkey's or subject's head and the necessary devices for the experiment in situ, with the calibration cube being positioned as close as possible to the eye. Thereby, crosstalk between the measurement channels, apparent crosstalk produced by magnetic fields, which are not exactly horizontal or vertical, and an approximation to the offset voltages are estimated",
            " To ensure simple interpretation of the measured data, the cube's axes must coincide with the coordinate system of the external stimuli. 2.2.2. Step 2: Calibration of the eye position with spontaneous eye movements. When the cube calibration is done, the subject or monkey performs spontaneous eye movements over the whole oculomotor range for 30- 60 sec. With these online recorded data, it is possible to calculate the gains of the measurement coils in the eye and the angle ~o between the frontal and the lateral coil (Fig. 1) by a numerical optimization procedure. With a similar procedure one can improve the values of the offset voltages. Finally, the straight ahead eye position has to be defined as reference position. For this, the subject or monkey has to fixate the zero-point once. 2.2.3. Step 3: Calibration with several fixation points. When the subject or monkey is able to fixate given targets with an error < 0.3 deg, it is possible to improve the calibration parameters with a modified optimization procedure. In this case, the subject has to fixate a number of given targets in the whole oculomotor range",
            " Quaternion representation of eye position was chosen because of its convenience for representing Listing's plane and to avoid misinterpretation of the torsional component which may occur in other coordinate systems, such as Fick or Helmholtz representation (\"false torsion\"; van Opstal, 1993). In order to measure the eye position in three dimensions, two tightly bound induction coils C1 (frontal coil) and C2 (lateral coil) are placed in two magnetic fields HH (horizontal) and Hv (vertical), which are quadrature modulated and spatially oriented approximately perpendicularly. By means of demodulation, the voltage induced by the horizontal field can be separated from that induced by the vertical field, giving at coil 1 the values Um and Ulv, and at coil 2 the values U2H and U2v (Fig. 1). The primary magnetic fields HI~ and Hv generate eddy currents in conductive surfaces, which in return generate magnetic fields He,. These generated fields superimpose on the primary fields. The quadrature demodulators detect the voltages which are in phase with an internal reference voltage (Fig. 1). According to the laws of induction and superimposi- tion the output voltage of one demodulator is Ud ----- gH' (C \"fill) -4- gv\" (C 'HH) -4- ~ ige,\" (C \"fie,) + 0 i (2) where: \u2022 c is the unit vector perpendicular to the surface of the recording coil; \u2022 o is the DC offset (voltages induced into stationary supply cables); \u2022 gH, gv, gei depend on the surface area of the measurement coil and the phase difference between each magnetic field and the reference voltage in the demodulator; \u2022 I--IH, fill, ~Iei are time-invariant vectors of field magnitudes pointing in the directions of the primary and secondary magnetic fields",
            " Given two unit vectors, cl and c'2, oriented perpendicularly to the planes of the two coils one gets the following output voltages: U1H = gin\" (C l\" HIH) -1- O111 (7) Ulv = glv\" (c1\" filv) + olv U2H = g211\" (C2\" H211) + O21t U2v = g 2 v (c2. fi2v) + 02v The indexes [-q~ff are omitted in these equations, but the introduced variables express effective gains and magnetic fields as described in Eqn (6). If the sensitivities g.., the offsetvoltages, the directions of the magnetic fields HH and Hv as well as the angle between both measurement coils tp are known, the vectors cl and c2 can be calculated directly from the measured voltages. 3.1.1. Calculation of ~ = Cly \u2022 Clz The coordinate system used here is defined in Fig. 1. The x-component of the vector Cl is always positive (one cannot look backward). Therefore, in order to calculate this vector, one has to solve the equation system c1\" Hl11 - gm | (8) cl Hlv glv | / C 1 = 1 such that Clx > 0. 3.1.2. Calculation of Cl = C2y \u2022 LCz~J The angle between both coils tp is regarded as constant. The cosine of this angle equals the scalar product of Cl and c2- As a result, one gets the linear equation system: . v_ vO2V / - l=k=cos / This equation system can be solved after calculating c~ from Eqn (8)",
            " The rotation matrix can now be transformed into other coordinates, such as quaternions, Fick- or Helmholtzcoordinates (Ferman, Collewijn, Jansen & Van den Berg, 1987; Tweed et al., 1990; Bartl, 1993). 3.2.1. Measuring field vectors using the calibration cube. The effective magnetic fields are measured by means of a special calibration cube, which contains three induction coils perpendicular to each other. In order to minimize the measurement inaccuracies caused by the field inhomogeneities, the calibration cube measurement must be performed as close as possible to the eye of the test subject (Fig. 1). The x-, y- and zoriented coils and the short circuit section of the cube are measured with each demodulator channel separately. When the first demodulator receives input from the short circuit section of the calibration cube, its output voltage is the offset associated with the demodulator (when driven by the calibration cube): U1Ho = O1Hcube When this same demodulator is driven, in turn, by the x-, y- and z-coils of the cube, its outputs are With x - coil : UIHx = alH \"Hmx + om~u With y - coil : U1Hy : alH \"Hmy + o 1 ~ (11) With z - coil : U1H z : am \u2022 HIH~ + O1H~be Computing the vector a 1H "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003766_0094-114x(95)00069-b-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003766_0094-114x(95)00069-b-Figure2-1.png",
        "caption": "Fig. 2. Geometric parameters of a serial kinematic chain. The calculation of the Jacobian needs the direction vectors e ~ of the joint axes, and the position vectors p~J of the joint axes with respect to the chosen reference frames and with respect to each other.",
        "texts": [
            " However, these indices would, by definition, always be identical, so the trailing ones are omitted. 3. J A C O B I A N M A T R I X FOR A S E R I A L C H A I N This section introduces the \"body-fixed,\" \"inertial\" and \"hybrid\" representations. It also repeats the well-known formal construction of the Jaeobian and its time derivative--in each of the representations--starting from the geometric parameters of the manipulator in a given position. Serial kinematic chains velocity mapping 139 3.1. Representations In general, a serial k inemat ic chain is given (at least) two reference frames, Fig. 2: (1) a \"ba se\" frame, {bs }, rigidly fixed to the world, i.e., the immobi le base of the chain, and (2) the \"end effector\" f rame, {ee}, rigidly fixed to the last link o f the chain. These two frames give rise to the following \" n a t u r a l \" representat ions o f the end effector twist: Body-fixed. In the body-f ixed representat ion, the end effector f rame {ee } serves also as reference f rame and its origin as the velocity reference point. Hence, the full nota t ion of a twist in this representat ion is ~ t ~e",
            " Since each co lumn of a Jacob ian matr ix J is a twist, the same set o f representat ions and representa t ion t r ans fo rmat ions exist for Jacobians. So, the symbol b,J j'b~ denotes the j t h co lumn o f J in the inertial representat ion. Similarly, e~J/'~ is the j t h co lumn in body-fixed representat ion, and b,J j'~ is the j t h co lumn in hybrid representat ion. I f an index is lacking, the corresponding representa t ion convent ion is assumed to be irrelevant. 3.2. Construction Consider a serial k inemat ic chain, with n revolute joints between base and end effector, Fig. 2. The base f rame {bs} is \" jo in t \" 0, and the end effector f rame {ee} is \" jo in t\" n + 1. Each co lumn of the Jacob ian matr ix is constructed based on the geomet ry o f the manipula tor . Let e ~ be the unit vector a long the ro ta t ion axis o f joint i, as well as one o f the three axes o f the or thogona l reference f rame {i} fixed to jo int i. Let p;J be the vector f rom the origin of {i} to the origin of {j}. Then, the columns j~.b, and J;'~ are given by [16]: 140 H. Bruyninckx and J"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure4-1.png",
        "caption": "Fig. 4. Expected pose of 3-R(RRR)R+R HAM in azimuth motion.",
        "texts": [
            " In the initial pose, a certain vector T V is given in the end coordinate system Tx T Ty and the direction of the vector is along the intersection line of plane OCT and plane Tx T Ty . The direction of vector T V is always along the intersection of plane OCT and plane Tx T Ty , and the rotation angle of the polarization mechanism is \u03b86 = 0\u25e6. Therefore, in the process of pitching, not only the antenna reflector does not move with it but also the polarization mechanism remains stationary. In the course of azimuth motion, the antenna pitch angle remains unchanged. As shown in Fig. 3, the expected pose of pitch motion is selected as the initial pose of azimuth motion and Fig. 4 is set as the desired pose of azimuth motion. In the course of azimuth motion, the HAM changes from the initial pose to the expected pose and the azimuth changes from \u03c2 to \u03c21. To keep the direction of vector T V always along the intersection of plane OCT and plane Tx T Ty , the rotation angle of the polarization mechanism is \u03b86. Two sets of transition poses can be used to realize the expected pose of the HAM from the initial pose of the azimuth motion to the expected pose, as shown in Fig. 5. https://doi"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000434_1.3643950-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000434_1.3643950-Figure3-1.png",
        "caption": "Fig. 3 Absolute va lue of dimensionless first-harmonic response (kxo,/Fi) a s a function of ratio of exciting frequency to first free-free beam frequency (oj/co1) for cases where dimensionless spring characteristic [a 2 = (\u00ab l)2/(fc/m)] is 10, characteristic force (Fi/fc) is 10 in. , with nonlinearity factor (e) a s a parameter",
        "texts": [
            " When the displacement amplitude of the first harmonic approaches asymptotically the natural frequency corresponding to the linear system supported by a rigid center support, the third harmonic of the force in the nonlinear spring approaches infinity which leads to infinite third-harmonic amplitude. This phenomenon accounts for peak 3. Peaks 4 and 5 both occur at a frequency equal to one third of the third natural frequency associated with the linear system resting on a linear spring with spring constant k. It may be seen from Fig. 3 that the force excitation for peaks 4 and 5 is triple-valued thus leading to three response curves for the third harmonic, the lowest response being too small to appear in Fig. 4. Journal of Applied Mechanics J U N E 1 9 6 0 / 2 7 3 Downloaded From: https://appliedmechanics.asmedigitalcollection.asme.org on 12/11/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use The reader should note the two different types of third-harmonic ''resonances\": 1 A linear system resonance excited by a finite third harmonic of the force in the spring (peaks 1, 2, 4, and 5)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003907_nme.1620380611-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003907_nme.1620380611-Figure7-1.png",
        "caption": "Figure 7. Conical helical bar with both ends fixed and under uniform loading",
        "texts": [
            " FIELD EQUATIONS OF THE NON-CIRCULAR HELICOIDAL BAR AND THEIR SOLUTION BY THE COMPLEMENTARY FUNCTIONS METHOD (CFM) The parametric equation of a helix is x = acos4, y = asin&, z = h4 The infinitesimal length element of the helix is defined as ds = (a2 + h2)1\u20192 d+ = c d 4 The relationships between the moving axes and the fixed reference frame are (Figure 2) {t,n,b}T = CB1 {i,j,k}T (254 where - (a/c) sin 4 (a/c) cos 4 - sin 4 (h/c) sin + - (h/c) cos 4 a/c 1038 V. HAKTANIR The curvatures of the circular helix are then given as x = a/cz, z = h/c2 The Frenet formulae for the helix are dt/d$ = (a/c)n, dn/d4 = (h/c)b - (a/c)t, db/d+ = - (h/c)n (27) The curvatures of the non-circular helicoidal bars are not constant along the axes. For example, radius of the helicoid for conical helix is given by (Figure 7) Variation of radius along the axes of hyperboloidal or barrel helix can be found in Reference 12. The rigidities may also be variable along the axes. For non-circular helicoidal bars under isothermal conditions, assuming (1) the (n, b) axes become the principle axes, (2) the warping is neglected and (3) initial twist of the cross-section is not considered, the static equations given in equations (21) and (22) will generalize to SPATIAL BARS 1039 These make up a set of 12 simultaneous differential equations with variable coefficients each one involving first degree derivatives with respect to position",
            " The boundary conditions, the loading and the material properties of the helix are as same as in the second example above expect the cross-section is circular. At the fixed ends the radius of the cross-section is ro = 15 cm and at the midspan rl . A parametric study is performed for different a = r l / ro as in Reference 25. The results obtained for this example are given in Table V and presented graphically in Figure 6. The displacements and rotations at the midspan (4 = 71/2) are presented in Table VI also. Example V. Conical helix with a constant circular cross-section. The conical helix is illustrated in Figure 7. The boundary conditions, the material properties and the loading are the same as in the previous examples. Circular cross-section is constant along the span and r = 15 cm, base circle has a radius of Ro = 310 cm. Using the notation given in Figure 7, a ratio fl = R1/Ro is used in parametric study. The force and moment components at the fixed ends are presented in Table VII also. The results are given in Figure 8. Example VZ. Conical helix with both endsJixed and with a linearly varying circular cross-section having one intermediate universal joint support under uniformly distributed vertical loading. The loading and the material properties of the helix are the same as in the previous example. The total angle of the system is n. At section 4 = n/2, there exists one universal joint support"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000650_978-3-319-42417-0_58-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000650_978-3-319-42417-0_58-Figure1-1.png",
        "caption": "Fig. 1. Quadruped robot model.",
        "texts": [
            " These facts suggest that head movements strongly affect the interlimb coordination with an increase in the locomotion speed. Owing to our effort in constructing a quadruped model using a simple CPG [5] with an additional postural reflex mechanism, we successfully reproduced gait transition from walk to pace to rotary gallop by using the head movement. Surprisingly, the obtained gait patterns are qualitatively similar to a giraffe\u2019s gait patterns [9], suggesting that these animals effectively make use of head and neck movements during gait transitions. Figure 1 shows an overview of the quadruped robot model constructed in this study. To mainly focus on the effects of the head movement on interlimb coordination, we used a simple musculoskeletal structure of the quadruped robot, which consists of a head, trunk, and four legs. The head and trunk segments are connected via a passive spring that has one degree of freedom (DOF) in the pitch direction only. We implemented two actuators for each leg (2 DOF), in which the rotational actuator drives a leg at the shoulder/hip joint and the linear actuator drives a leg along the leg axial direction, as shown in Fig. 1 right. A phase oscillator was implemented in a leg to generate a rhythmic leg motion during the stance and swing phases according to the oscillator phase \u03c6i (hereafter, i = 1 \u223c 4, Left fore (LF):1, Left hind (LH):2, Right fore (RF):3, Right hind (RH):4), as explained in Sect. 2.2 in detail. We implemented two types of sensors: pressure sensors on the feet, which can detect ground reaction forces Ni [N] (GRFs) parallel to the leg axis, and an angle sensor on the neck joint, which can detect the displacement \u03c8 [rad] from the equilibrium posture",
            " The target angle \u03b8\u0304i [rad] and length l\u0304i [m] are described using the oscillator phase as follows: \u03b8\u0304i = \u2212Camp cos \u03c6i, (1) l\u0304i = L0 \u2212 Lamp sin \u03c6i, (2) where Camp denotes the amplitude of the target angle, and Lamp and L0 denote the amplitude and offset length of the target length, respectively. Based on the target angle and length, each actuator is controlled using the following equations: \u03c4i = \u2212Kr(\u03b8i \u2212 \u03b8\u0304i), (3) Fi = \u2212Kl(li \u2212 l\u0304i), (4) where Kr and Kl are the P gains for each actuator. \u03b8i and li denote the actual angle of the shoulder/hip joint and the actual leg length, respectively. Based on the control scheme, a leg tends to be in the swing phase for 0 < \u03c6i \u2264 \u03c0, and in the stance phase for \u03c0 < \u03c6i \u2264 2\u03c0, as shown in Fig. 1 right. The dynamics of the phase oscillators are described as follows: \u03c6\u0307f i = \u03c9 \u2212 \u03c3Nf i cos \u03c6f i \u2212 \u03c1\u03c8 cos \u03c6f i (5) \u03c6\u0307h i = \u03c9 \u2212 \u03c3Nh i cos \u03c6h i + \u03c1\u03c8 cos \u03c6h i (6) where \u03c6f i and \u03c6h i denote the phases of the fore and hind leg oscillators, respectively. \u03c9 [rad/s] denotes the intrinsic angular velocity of the oscillators. \u03c3 [rad/Ns] and \u03c1 [1/s] denote the feedback gains for the second and third terms, respectively. Nf i and Nh i [N] denote the GRFs detected by pressure sensors in the fore and hind feet, and \u03c8 [rad] denotes the neck joint angle detected by the angle sensor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000869_s11668-019-00593-2-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000869_s11668-019-00593-2-Figure3-1.png",
        "caption": "Fig. 3 Schematic of a pinch roll bearing used in HSM [7]",
        "texts": [
            " These are driven by a motorconnected bearing assembly. Failure occurred in the bottom entry pinch roll in its drive side as shown in Fig. 2c. The bearing was found to be jammed and broken into multiple pieces. It failed after 5 months of service as against an expected life of at least 1 year. This reduction in service to less than half of its expected life is the motivation behind an in-depth failure analysis to understand its root cause and take corrective actions to prevent such failures in the future. Figure 3 shows a schematic of a pinch roll bearing used in HSM. It has three major elements, namely outer ring, inner ring and barrel-shaped rollers. There are two sets of rollers partitioned by a separator as shown in Fig. 3. The mechanism of operation of the bearing is such that rollers roll around the inner race like a car wheel, but traverse around the circumference much slower than the inner race. The outer ring is stationary, whereas the inner ring revolves which drives the shaft. There was an important finding during the on-site observation. It was observed that rollers on the near end of the actual roll were broken and were found to be in severely damaged condition, whereas those located toward the near end of the shaft remained in an undamaged state"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001473_rpj-04-2019-0113-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001473_rpj-04-2019-0113-Figure3-1.png",
        "caption": "Figure 3 Schematic diagrams of compression tests",
        "texts": [
            " Each value of microhardness was the average of Table I Chemical compositions of Mg alloy and ZM5 (Wt.%) Al Zn Mn Mg AZ61D Mg alloy 6 1 0.5 Bal. ZM5 8.5 0.5 0.2 Bal. Effect of laser power Xingcheng Wang, Changjun Chen and Min Zhang Rapid Prototyping Journal Volume 26 \u00b7 Number 5 \u00b7 2020 \u00b7 841\u2013854 ten points measured in a random way. Compression specimens were tested by electro-mechanical universal testing machine (RGM-4100), with the compression speed set as 1mm/min. The compression test diagram is shown in Figure 3. The sizes of horizontal and vertical compression specimens are 12 6 6 and 6 6 12 mm3, respectively. Subsequently, the fracture morphology was characterized by SEM. Figure 4 presents the surface morphology of SLMed samples at different laser power. It is obvious that the surface morphology ofMg\u2013Al\u2013Zn alloys built by SLM is affected by the laser energy input. As shown in Figure 4(e) and (f), when the laser power above 100W, samples surfaces are rough and uneven, crisscross and deep gully at the center of laser scanning tracks (CLST) are induced"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000204_j.autcon.2019.102996-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000204_j.autcon.2019.102996-Figure3-1.png",
        "caption": "Fig. 3. Notch outline curves; distribution on a hub cross-section; examples of different types of hub profiles.",
        "texts": [
            " The first is a global definition of a discretized surface (mesh), support locations, and angles between elements. The second is a local definition of the connection geometry, which takes the global mesh as input. The local and global models are closely related, and local definition inputs should be compatible with the global definition. Compatibility settings include parameters such as the scale of the interconnecting parts and hubs, the size, quantity and configuration of screws, and the number of notches allowed in each type of hub. Fig. 3 shows examples of notch outline curves, of distribution around the edge of the hub's cross-section, and of types of hubs with different crosssections, notch geometry, and notch distribution. The choices of best hub or shape proportions are not in the scope of this work. Once the global and local settings are compatible, the generation of the geometry of the interconnecting parts is automatic. Fig. 4 presents the concept of the connection system for a single and a double layer gridshell. A single layer connection system is in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.14-1.png",
        "caption": "Fig. 10.14 Structure of reciprocating disintegrator",
        "texts": [
            " If the packed refined cellulose is used for nitration, disintegration is required before use to ensure the refined cellulose is fluffy, which is conducive to secondary air drying and nitration. Continuous nitration is normally combined with the continuous disintegrators, which are classified into horizontal and reciprocating disintegrators; horizontal disintegrator has a low disintegration efficiency and affords a poor quality uniformity. Before the end of disintegrating each pack of cellulose, the whole piece of fiber will be transferred into the disintegrator [2]. The structure of the reciprocating disintegrator is shown in Fig. 10.14. Reciprocating disintegrator not only can move forward and back forward but also can shift in the process of reciprocating to ensure the uniformity and consistency of feeding. 10.2.3.3 Bleaching Process of Refined Cellulose For the production of NC with low viscosity, in addition to continuous digestion that is used to reduce the viscosity of NC, the bleaching process can be used to reduce further and control the viscosity of NC. For a special industrial application, to achieve a requirement of whiteness, cellulose must be bleached in the refining"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003709_s0043-1648(96)07486-8-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003709_s0043-1648(96)07486-8-Figure7-1.png",
        "caption": "Fig. 7. Zero-sliding points in a spherical roller thrust bearing.",
        "texts": [
            " If sliding occurs over the whole slice, the following expression for Coulomb friction is used: tsmp (10) Due to the curved contact surface in a spherical roller thrust bearing, the rollers will undergo sliding in the contact. For a roller there will be two points along each contact where the sliding velocity is zero. These zero-sliding points form the generatrices of a rolling cone, which represents the surface on which pure rolling occurs. Sliding is present at all other points along the contact, in the direction of rolling or opposite to it, depending on whether the roller radius is greater or smaller than the radius to the rolling cone (see Fig. 7). For an unskewed roller, the generatrices of the rolling cone and the roller middle line intercept the bearing centre axis at the same point. Another assumption for the roller to be in equilibrium is that the total tangential load is zero for each contact (the roller\u2013housing washer contact and the roller\u2013 shaft washer contact). The two zero-sliding points are determined under these two assumptions, minimising the tangential load in an iterative manner. Given the two zero-sliding points, the creep ratio, j, can be calculated by determining the roller radius, R r , and the radius to the rolling cone, R c , at each disk: R yR r c js (11)* *R r Two test cases were used to study the roller\u2013housing washer contact, with different values for the coefficient of friction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002107_jestpe.2020.3048091-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002107_jestpe.2020.3048091-Figure5-1.png",
        "caption": "Fig. 5. Schematic diagram of PMSM controlled by RRCC.",
        "texts": [
            " In the first condition, one voltage vector is not enough to cover the distance, so the oriented unit voltage vector is first gained from the fraction and then amplified by um. In the second condition, if the magnitude is enough, the voltage drop on the resistance will be compensated. In the other conditions, us(k) can be output directly. That is the whole process of the proposed RRCC in one cycle, and the flowchart is presented in Fig. 4. The other peripheral parts of PMSM control with RRCC are illustrated in Fig. 5. The main feature of RRCC is to utilize a voltage vector sequence to aim at one target, the terminal point of a future stator flux linkage vector reference. A single voltage vector can be used to either lengthen the magnitude of stator flux linkage vector or increase the torque angle, or both. However, the longer the stator flux linkage vector is, the smaller the change in torque angle could be. That means torque must vary slowly with long stator flux linkage vector. In most torque control strategy, both the length of stator flux linkage vector and the torque angle increase and decrease synchronously during the torque tracking process",
            " In RRCC, the main part of the trajectory looks straight as shown in Fig. 2. Shorter trajectory of less voltage vectors means less cycles and faster torque dynamic response. Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 19,2021 at 05:02:35 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (c) 2020 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. Fig. 5 could be seen as a general schematic diagram of PMSM control using the state equations in FOC framework. All the parameters of the tested PMSM, inverter and DC voltage source can be used directly. Thus, the block of \u201cProposed RRCC\u201d can be replaced with other torque control strategy. The simple schematic diagrams of PI-based FOC (with feedforward items), MPTC and MPCC are presented in Fig. 6. Then, the main process of RRCC could be discussed to compare with other strategies. The main process of RRCC contains three steps"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002221_0278364919897134-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002221_0278364919897134-Figure12-1.png",
        "caption": "Fig. 12. Exploded view of membrane clamping assembly. Compressed air can be released under the membrane to inflate it once it is clamped down. Twelve screws and a clamp ring provide an even clamping force on the membrane. Stainless steel dowel pins provide anchor points for the radial fibers and help to align the membrane on the clamp base.",
        "texts": [
            " We can then characterize the polar angle f and its spatial derivative as f = sin 1 r R , \u00f08a\u00de df dr = 1 R ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 r R 2 q \u00f08b\u00de As shown in (5), the magnitude of the differential fiber length with respect to the parameter g can be defined by its final configuration. This corresponds to when the fibers form major arcs along the f direction. Presenting this mathematically, we obtain dL dr = R df dr f = 1ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 r a 2 q \u00f09\u00de where a is the inner radius of the membrane clamp ring (shown in Figure 12), which is the radius of the final hemisphere, R. Finally, the unknown component of the reorienting fiber can be solved for by applying the relations established in (8b) and (9) to (7) and rearranging the expression to give du dr = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 (r a ) 2 1 1 (r R ) 2 q r \u00f010\u00de Then, solving for u as a function of parameter r simply requires integration, u = u0 + a r ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 r a 2 r ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 r R 2 r ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 1 r a 2 1 1 r R 2 s F sin 1 r a , a R 2 \u00f011\u00de where F(c,m) is the incomplete elliptical integral of the first kind, see Abramowitz and Stegun (1972)",
            " As the applied pressure was increased by adjusting the regulator, a 20\u2013250 kPa absolute pressure sensor (MPXHZ6250AC6T1, Freescale Semiconductor) sampled by a 16-bit ADC (MAX1167 BEEE + , Maxim Integrated Products) controlled by an Arduino Mega 2560 provided an average value for the pressure over a 1 second interval. While the pressure was averaged, pictures were taken of the membrane from the front (to show the fiber pattern) and side (to calculate volume) using 8 MP digital cameras (Figure 11). The membrane was fixed to our setup for pressure testing using the clamping assembly shown in Figure 12. A quick connect tube fitting for 8 mm tube was threaded into a 0.25 in thick laser-cut (ULS PLS6MW, 50 W CO2 laser) delrin base. Blind holes were drilled into the base for dowel pins, which were placed at the end of each fiber to prevent the fiber from pulling the membrane out of the clamp. This also ensured that the fibers maintained the correct orientation during clamping and testing. A 3.2 mm thick ring with an inner diameter of 9.37 cm was used to clamp down on the edge of the membrane. Helicoil inserts in the base mated to 12 4\u201340 screws to ensure a strong and even clamping force"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001438_lra.2020.2969161-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001438_lra.2020.2969161-Figure1-1.png",
        "caption": "Fig. 1. The prototyped microsurgical forceps-driver and master device.",
        "texts": [
            " Also, the function to provide the haptic feedback on the gripping-force was not implemented on the master device, so the surgeon could not perceive whether the gripping-force applied to the tissue is proper. In this work, considering the essential functions that must be implemented in the forceps-driver and the master device of the tele-operated microsurgical robotic system, we aim to develop a novel sensor-embedded forceps-driver with high-precision gripping-force control and a master device with the function to display haptic feedback on the gripping-force. The conceptually designed both devices are fabricated as functional prototypes (see Fig. 1), and the working performances are experimentally investigated. Also, the practical feasibility and applicability are assessed through the user test performing the tele-operation utilizing the proposed system. The structural configuration and the working principle of the forceps-driver are schematically shown in Fig. 2(a) and (b-i), respectively. In designing the forceps-driver, we focus on implementing the device having the functions to drive commercial microsurgical forceps and to measure the gripping-force applied to the gripped-object by the forceps-tips"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001298_tmag.2019.2942023-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001298_tmag.2019.2942023-Figure2-1.png",
        "caption": "Fig. 2. Coordinate transformations between the stator reference frame and the rotor reference frame.",
        "texts": [
            " In this article, the first-order perturbation is considered only, which means that all the quantities of the magnetic fields, such as the magnetic vector potential Az(r, \u03b1, \u03b5), the radial component of the magnetic flux density vector Br (r, \u03b1, \u03b5), and the tangential component of the magnetic flux density vector B\u03b1(r, \u03b1, \u03b5) are the linear function that depends on the eccentricity \u03b5 [28] Azi (r, \u03b1, \u03b5) = A0 zi (r, \u03b8)+ \u03b5A1 zi(r, \u03b1)+ \u00b7 \u00b7 \u00b7 (1) Bir (r, \u03b1, \u03b5) = B0 ir (r, \u03b1)+ \u03b5B1 ir (r, \u03b1)+ \u00b7 \u00b7 \u00b7 (2) Bi\u03b1(r, \u03b1, \u03b5) = B0 i\u03b1(r, \u03b1)+ \u03b5B1 i\u03b1(r, \u03b1)+ \u00b7 \u00b7 \u00b7 (3) where the superscripts 0 and 1 indicate the zeroth-order and first-order quantities, the subscript i denotes the sub-region for analysis, and r and \u03b1are the radial and tangential components, respectively. B. Coordinate Transformations Fig. 2 shows the coordinate transformations between the stator reference frame r \u2212\u03b1 and the rotor reference frame \u03c1 \u2212 \u03c8 . The transformation relation can be denoted as follows [29]: \u03c1 = r \u2212 \u03b5 cos(\u03b1 \u2212 \u03c6)+ O(\u03b52) (4) \u03c8 = \u03b1 \u2212 \u03c9t + \u03b5 sin(\u03b1 \u2212 \u03c6)/r + O(\u03b52) (5) where \u03b5 and \u03c6 can locate the position of the rotor. The first-order component is retained, which can meet the precision demand of the magnetic field. III. FIELD SOLUTION In this article, the solution of the magnetic field has been derived. The governing equations in each subdomain are listed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000945_tcyb.2019.2903852-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000945_tcyb.2019.2903852-Figure1-1.png",
        "caption": "Fig. 1. Examples of the constraint operator.",
        "texts": [
            " Definition 1 [20]: Defining SQi(\u00b7) is a constraint operator such that SQi(x) = (x/||x||) max0\u2264 \u2264||x||{ |(\u03b1 x/||x||) \u2208 Qi, \u22000 \u2264 \u03b1 \u2264 1} and SQi(0) = 0, where Qi \u2286 R l, i = 1, . . . , n, is a nonempty bounded closed nonconvex set and 0 \u2208 Qi. Moreover, maxx\u2208Qi{\u2016SQi(x)\u2016} = \u03bc\u0304i > 0, infx/\u2208Qi{\u2016SQi(x)\u2016} = \u03bc i > 0, where \u03bc\u0304i and \u03bc i are two positive constants. Remark 1: The function of the operator SQi(x) is to obtain the vector with the largest magnitude whose direction is the same as x, \u2016SQi(x)\u2016 \u2264 \u2016x\u2016 and SQi(x) \u2208 Qi for all 0 \u2264 \u2264 1. Some examples of this constraint operator are shown in Fig. 1. Note that there is no convexity assumption on each Qi. maxx\u2208Qi{\u2016SQi(x)\u2016} = \u03bc\u0304i > 0 means that x is bounded and infx/\u2208Qi{\u2016SQi(x)\u2016} = \u03bc i > 0 means that x can be in an arbitrary direction. The dynamics of each second-order follower i \u2208 F with the nonconvex velocity constraint is xi((k + 1)T) = xi(kT) + vi(kT)T vi((k + 1)T) = SQi [vi(kT) + ui(kT)T] (1) where xi \u2208 R l, vi \u2208 R l, ui \u2208 R l, and Qi are the state, the velocity, the input, and the velocity constraint of the i follower, respectively. Note that R l is the set of real column vectors of the l dimension"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001975_10402004.2020.1836294-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001975_10402004.2020.1836294-Figure1-1.png",
        "caption": "Fig. 1. Structure of the first generation gas foil bearing.",
        "texts": [
            " Different stiffness matrices with respect to frictional contacts are added to the FE equilibrium equation depending on the contact states at each incremental step. To obtain the pressure distribution, the Reynolds equation (RE) is solved using the finite difference (FD) method and then coupled with the structure domain. The model is efficient in terms of computational time. Acc ep te d M a us cr ipt The structure of the first generation GFB mainly consists of two parts, i.e., a smooth top foil and an underneath corrugate bump foil, as shown in Fig. 1. Both the top foil and bump foil are fixed to the rigid sleeve at one end by spot welds, while free at the other end. In the bearing system, hydrodynamic pressure occurs within the fluid film between the top foil and the shaft surface with the rotation and eccentricity of the shaft. The bump foil is in contact with the top foil and the bearing sleeve. These frictional contacts result in structural damping and energy dissipation during operation, exerting significant influences on the performance of GFBs",
            " For the present model and the NDOF model, the deflections of the first two bumps are negative under the increasing/decreasing load. This is because they are pushed by the latter bumps toward the fixed end. Pressure Distribution and Film Thickness Acc ep te d M an us cr ipt The pressure distribution of the GFB with a rotational speed of 30000 rpm and a load of 134.1 N is depicted in Fig. 7 (a). It is noted that the shaft rotates from the free end to the fixed end of the foil structure, as shown in Fig. 1. It can be seen in Fig. 7 (a) that the region of high pressure is narrow in the circumferential direction; hence, the pressure is mainly applied to the several bumps in this active region while the bumps near the fixed end and the free end receive much less load. Near the fixed end, the pressure is the same as the ambient pressure because the G\u00fcmbel boundary condition is applied. Fig. 7 (b) depicts the vertical deflection of the top foil in the circumferential direction under the pressure. It can be seen that the predicted deflection has a wavy appearance because the top foil between adjacent bumps tends to sag under pressure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000668_tcst.2016.2601286-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000668_tcst.2016.2601286-Figure3-1.png",
        "caption": "Fig. 3. Exploded view of a split gear assembly.",
        "texts": [
            " As the changes in the gear contact torque direction are inevitable when applying a constant generator torque, measures have to be taken to improve the range extender unit\u2019s Noise Vibration Harshness behavior. Split gears are the wellknown devices to avoid backlash or play in gearing applications where smooth operation in starting conditions and while reversing the direction of motion is required [6]. In addition to backlash-free positioning, split gears have recently been used for higher power applications, e.g., engine camshaft drives [7]. In the following, a custom-made split gear is evaluated as a countermeasure to avoid rattling. Fig. 3 shows a drawing of a split gear assembly, consisting of two axially divided half gears. In between the two halves, there are \u201comega\u201d springs, which are loaded if the halves are rotated against each other. In one direction, the gear half which is directly mounted to the shaft transmits the torque. In the other direction, the torque goes from the floating gear half over the spring to the hub and subsequently the shaft. Therefore, the preload torque has to be higher than the highest torque transmitted over the floating gear half to avoid gear rattle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002822_j.aime.2021.100040-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002822_j.aime.2021.100040-Figure2-1.png",
        "caption": "Fig. 2. Considered half of the build chamber for simulation, plane of symmetry marked in blue.",
        "texts": [
            " A prerequisite is that the monosilane content is higher than the residual oxygen and moisture content. For the calculations an atmospheric pressure of 101,3 kPa and an atmospheric temperature of 20 C were assumed. The mass of formed SiO2 was modeled for different residual oxygen and moisture contents as well as for different test chamber volumes. Based on the general requirements established in phase 1, a simplified build chamber was designed using the CAD (computer-aided design) software SolidWorks (Dassault Syst emes). As it can be seen in Fig. 2, the chamber has two gas inlets and one gas outlet. The main gas inlet at the bottom is used to create a gas stream above the powder bed to remove process emissions while the secondary inlet serves to prevent contamination of the protective glass by particles or smoke. Recoater, powder supply and building platform, that are essential parts normally found in the build chamber, were neglected for the simulation. This is because the recoaters position and the powder supply are outside the gas flow during processing and the influence of the building platform is expected to be small in comparison with the influence of the inlet design. Additionally, since the regarded geometry of the build chamber is symmetrical, it was sufficient to simulate only the flow within one half of it (Fig. 2). This way, the calculation effort and time could be reduced. In order to prepare the geometry for the fluid simulation, the opensource software Salome was used for pre-processing. The open-source CFD (computational fluid dynamics) toolbox OpenFOAM\u00ae (Open Field Operation and Manipulation) was employed to conduct the CFD simulation of the gas flow within the build chamber. To discretize the simulated fluid volume a simple base mesh was generated. This base mesh was adapted to the simulated geometry and refined with SnappyHexMesh"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002162_978-3-030-48977-9_1-Figure1.5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002162_978-3-030-48977-9_1-Figure1.5-1.png",
        "caption": "Fig. 1.5 Example of commercial converter unit and integrated semiconductor module which can be used to build an inverter [16]",
        "texts": [
            " \u2022 Availability of compact high-performance digital processors with extensive I/O capabilities that can be readily interfaced with the equally compact power electronic drive circuitry required to control the switching devices. \u2022 Design and manufacturing improvements in passive devices most notably in capacitors which play a key role in terms of overall voltage source converter sizing and costs. The culmination of the improvements indicated above is exemplified by the availability of building blocks as shown in Fig. 1.5, to construct a complete inverter and \u201coff-the-shelf\u201d commercial inverters, as shown in Fig. 1.6, which can be readily interfaced to electrical machines. Note that the term inverter refers to a dc to ac converter, as shown in Fig. 1.1. In this section, the emphasis has been predominantly placed on improvements in volumetric power density. However, improvements in converter technology leading to the ability to operate at much higher electrical fundamental frequencies have also been instrumental in realizing high speed drives"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002085_tec.2020.3044000-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002085_tec.2020.3044000-Figure9-1.png",
        "caption": "Fig. 9. No-load magnetic flux lines for (a) 39-slot and (b) 36-slot with 12-pole motors",
        "texts": [
            " Downloaded on December 18,2020 at 20:38:30 UTC from IEEE Xplore. Restrictions apply. 0885-8969 (c) 2020 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. 8 the motors keeping Ampere x turns the same for both the integral-slot balanced and the fractional-slot UBW motors. The rotor and stator of both machines are un-skewed. No-load magnetic flux lines for 39-slot and 36-slot with 12-pole motors are illustrated in Fig. 9. No-load line back-EMF voltages of 39-slot with UBW machine (UBWM) and 36-slot with ISW machine (ISWM) are obtained for various magnet pole-arc values (Fig. 10). It can be seen that THD of line back-EMF voltage for 39-slot with UBW machine is always lower than that of conventional 36- slot ISW machine. Furthermore, the minimum THD value for UBW motor is achieved as 0.26 % while it is 2.94 % for ISW motor at the same magnet pitch, which demonstrates one of the main benefits of using such a winding structure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000524_s0005117916010069-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000524_s0005117916010069-Figure1-1.png",
        "caption": "Fig. 1. Tracking geometry on a plane. E0,P0 and Et,Pt are initial and current locations of target and observer.",
        "texts": [
            " The auxiliary problem is to get the most accurate TME estimate with noisy current bearing only measurements. The criterion is a certain terminal (\u201ccompromise\u201d) payoff functional that reduces two-criterial problem to a single-criterial one. This problem setting is relevant for underwater vehicles. Consider a motion of both the observer P and the target E in a fixed Cartesian coordinate system Oxy whose origin O coincides with the observer\u2019s location at the initial time moment t0 = 0, and the Oy axis is oriented towards the first bearing measurement (Fig. 1). TMEs are characterized by a vector of parameters \u03b8 = (\u03b81, \u03b82, \u03b83), where \u03b81 and \u03b82 are the components of target\u2019s velocity vector along axes Ox and Oy respectively, and \u03b83 is a distance from P to E at the initial time moment. The observer P on each tick of the motion t = 0, 1, . . . , T produces a measurement \u03bet of the bearing \u03d5t to the target E with some additive noise \u03c3 \u03b5t, where ( \u03b5t ) t=0,1,2,... is a sequence of independent standard Gaussian random values, \u03c3 = const > 0 (mean squared bearing error)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001811_s12008-020-00670-z-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001811_s12008-020-00670-z-Figure8-1.png",
        "caption": "Fig. 8 3D manipulator models used to explain the DH method using computer simulations",
        "texts": [
            "\u201d In addition, the professor conducted an exam to evaluate the student\u2019s ability to apply the learned knowledge to solve a real problem. The steps of phases 2 and 3 of this study are described in detail below: 1. Theoretical explanation of the DH method using computer simulations: The professor explained the steps and rules of the DH method using simulations of the 3D manipulators with the experimental platform. The explanation began with a step-by-step demonstration of how to apply the DH method in a 1-DoF serial manipulator. The explanation was repeated using manipulators with greater DoFs (see Fig.\u00a08). Two of the key concepts analyzed for each manipulator were the definition of the base reference frame (also called world reference frame) and the possible options to define the TCP. This step finished when the students had no questions about how to apply the rules of the DH method. 2. Explanation of the key characteristics: Using computer simulations, the professor explained key characteristics that students must consider in order to solve a problem correctly, considering the conditions of the DH method"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001347_012014-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001347_012014-Figure1-1.png",
        "caption": "Figure 1. Schematic diagram of the first generation PFM with differentiated friction pairs.",
        "texts": [
            " These conditions reflect the dependences of the magnitude of the coefficient of force and the spacer force generated by the feedback control device on the current coefficient of the gain value of the coefficient of friction, which theoretically ensure the limiting value of the torque transmitted by the clutch, which leads to failure-free operation of the crank presses. A schematic diagram of a first-generation SFC with single-circuit negative feedback having differentiated friction pairs of the type \u201cleading pairs - driven pairs\u201d is shown in Fig. 1 (the upper part of DTS-2019 IOP Conf. Series: Materials Science and Engineering 680 (2019) 012014 IOP Publishing doi:10.1088/1757-899X/680/1/012014 the figure relative to the axis of rotation of the coupling). Two kinematic half-couplings 1 and 2 coaxial with each other are connected to each other in the circumferential direction of the friction group consisting of friction disks 3 and 4. The disks 3 are connected in the circumferential direction with the hub of the pressure disk 5, the disks 4 are connected to the half-coupling drum 2. The pressure disk 5 is devoid of kinematic connection with the hub of the coupling half 1 in the circumferential direction, with the exception of slight friction between them, which is not considered further [1-3]. The control device is made in the form of rolling bodies 6, which are placed in beveled sockets made on the end surfaces of the pressure disk facing one another and rigidly fixed on the hub of the coupling half 1 of the thrust disk 7 (Fig. 1, section A-A). The force closure of friction pairs is created by a spring 8, the force of which is transmitted to the pressure disk through a thrust bearing 9 to reduce friction between them[4-6]. Given the action of the spacer force in the SFC (see Fig. 1, section AA), we write the formula for the torque in the form: ( ),p p\u0422 zR f F F= \u2212 (1) F \u2212 spacer force; other designations are looked above. When the structural-layout scheme of the friction group, built according to the type of \"leading friction pairs driven friction pairs\", the spacer force is determined by the formula: ( )1 tgpz z \u0422 F zr \u2212 = \u03b1 , (2) 1z \u2212 the number of leading friction pairs of the friction group. Putting the expression (2) in the formula (1), we obtain: 1 , 1 ( )p p f\u0422 zF R z z Cf = + \u2212 (3) The greatest accuracy of the SFC operation will be when constp\u0422 = ",
            " The same conclusion can be drawn from the analysis of relation (8). Therefore, the implementation of the limitation of the magnitude of the gain at the maximum value of the coefficient of friction is possible only in the multi-disk version of the SFC with differentiated pairs of friction. The number of leading friction pairs should be minimal. We are also investigating a variant of the structural-layout scheme of the friction group in which friction pairs are not differentiated into leading and driven, but all are driven. The indicated option is shown in fig. 1 (lower part relative to the axis of rotation of the coupling) [16-19]. The difference between the considered SFC variant is the introduction of a thrust bearing 10, installed between the leftmost (in Fig. 1) friction disk 4 and the thrust disk 7. The kinematic connection between them is broken, and the entire load of the coupling is transferred from the coupling half to the coupling half 2 (or vice versa) by means of rolling 6. Applying the same method as for the SFC variant with differentiated friction pairs, the calculation gain method, we find the expression for calculating the coupling torque: . 1p p f\u0422 zF R zCf = + (9) Relation (9) differs from relation (3), with the same parameters, by a lower value of the reduced coefficient of friction equal to: , 1p ff zCf = + and, accordingly, lower torque p\u0422 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002048_0142331220966427-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002048_0142331220966427-Figure2-1.png",
        "caption": "Figure 2. The world and body coordinate system.",
        "texts": [
            " Section 2 briefly describes the dynamic model of the quadrotor with tiltable rotors. Section 3 introduces the ESO-based omnidirectional control and control allocation strategy. Section 4 analyzes the observer stability and cloesd-loop stability. Extensive numerical simulations are presented in Section 5, while conclusions and future directions will be explained in Section 6. Let FW : OW : XW ,YW ,ZWf g denote the world inertial frame fixed on the ground, and FB : OB : XB,YB,ZBf g denote the body frame attached to the vehicle body at the center of gravity, as seen in Figure 2. There are also four frames of the propeller groups FPi : OPi : XPi ,YPi ,ZPi f g shown in Figure 3. These frames rotate around their XPi axes, w.r.t the body frame. As usual, the rotation matrix WRB stands for the orientation of the vehicle body frame w.r.t world inertial frame while BRPi stands for the orientation of the i-th propeller group frame w.r.t body frame. By denoting the propeller tilting angle around the axis XPi with ai, it follows that BRPi =RZ p 2 i+ p 4 RX (ai), i= 1 . . . 4: \u00f01\u00de Also let OB Pi =RZ p 2 i+ p 4 l 0 0 2 4 3 5, \u00f02\u00de be the origin of FPi in FB with l being the distance between OPi and OB"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003023_s11071-021-06591-0-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003023_s11071-021-06591-0-Figure1-1.png",
        "caption": "Fig. 1 Simplified overlooking diagram of robotic fish",
        "texts": [
            " 3, the data-driven dynamic model of the RF considering the influence of head-yawing is established. In Sect. 4, the proposed DITSMC-WHY based on the dynamic model is designed. In Sect. 5, real-time experiments on the RF platform are carried out to demonstrate the efficacy of the propose control. Section 6 gives the conclusion. Based on the biological model of bluefin tuna, whose posterior 1/3 to 1/2 provides forward power by oscillation [16, 17], the simplified RF overlooking diagram is shown in Fig. 1. The coordinate origin is established at the junction of the fish head and the body, which is also the center of gravity of the RF. The direction of the fish tail is the positive direction of the x-axis, and the positive direction of the y-axis can be defined according to the Cartesian Coordinate system, which is defined as the longitudinal position of the fish body at t moment on the x-axis. Thus, the continuoustime fish body wave of the RF can be written by [13]: y x; t\u00f0 \u00de \u00bc C1x\u00fe C2x 2 sin rx\u00fe xt\u00f0 \u00de \u00f01\u00de where y x; t\u00f0 \u00de is the longitudinal position of the RF body when swinging; x is the horizontal position of the RF body; t is the time; r \u00bc 2p=k is the fish body wave number, where k is wave length; x \u00bc 2pf \u00bc 2p=T is the swing angular velocity of the RF body, where f is the body wave frequency and T is the period of the wave; C1 and C2 are coefficients of envelope line for the fish body wave function, where C1 is the first-order coefficient of the wave envelope and C2 is the quadratic term coefficient of the wave envelope",
            " The information of the actual speed and the yawing angle is sent and demonstrated in the upper computer wirelessly. In order to better fit the body wave of the biological fish, the physical prototype of the RF is depicted in Fig. 4. The fusiform head is used to accommodate the battery, the DSP-based master control and wireless communication modules, mobile station, and gyroscope. The flexible fish body with four connected joints covered by a latex skin and the caudal fin are also designed. The physical mass (M) is 4:2 kg, while the lengths of the RF and fusiform head (d in Fig. 1) are 0:67 m and 0:34 m, respectively. The RF body links consist of DC motors as well as other supporting components. For fitting the tail of the real fish, the aluminum alloy skeleton is used in the RF caudal fin to connect to the fourth joint motor by using silica gel. In this paper, we select the RF oscillating frequency to be the system control input, which can be obtained by the controller based on the real-time error calculated by the reference forward speed and the actual speed at each moment"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000234_ecce.2019.8913191-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000234_ecce.2019.8913191-Figure6-1.png",
        "caption": "Fig. 6. Illustrations of concepts of PM arrangements and extended flux barriers to enhance magnetization characteristic.",
        "texts": [
            " On the other side, when VPMs are negatively magnetized, initial operating point is point-B2 shifted from point-B1. In order to change the magnetization ratio to around +90%, the operating point of VPMs is needed to attain pointB3. In general, positive magnetizing current is larger than negative one because of shifted load line by CPMs in case that VFMMs have parallel circuit configuration. However, designing positive and negative current relatively equally is important to reduce the inverter rated output capacity. III. CONCEPT OF ROTOR SHAPE IN THE PROPOSED VFMM Fig. 6(a) illustrates conventional VFMM having parallel circuit configuration. If VPMs are demagnetized, the closed magnetic circuit is made in the rotor. The closed magnetic circuit contributes to reduction of the iron loss and induced voltage. Meanwhile, in this PM arrangement, VPMs are affected by the diamagnetic field from CPMs as mentioned in previous section. As a result, the operating point of VPMs is largely shifted to negative side. Eventually, asymmetric positive and negative magnetizing current occur. RCPM RVPM FCPM FVPM CPM VPM gap Rgap Rstator+ FW inding Rrotor RCPMFCPM CPM RVPM FVPM VPM gap Rgap Rstator+ FW inding Rrotor (a) Series hybrid PMs. (b) Parallel hybrid PMs Fig. 4. Equivalent magnetic circuits of series and parallel flux paths. B. VFMM combining series with parallel circuits Fig. 6(b) shows novel rotor structure which possesses additional CPM near by surface of the rotor. The additional CPM have series relationship to VPMs, and consequently, the rotor structure has combining series with parallel circuit configuration. Outer CPMs can cancel the diamagnetic field against VPMs as shown in Fig. 6(b) because some magnetic fluxes generated by additional CPMs behave as the magnetic field to strengthen the positive magnetization state of VPMs. Accordingly, this rotor structure using double layer CPMs has potential to weaken the diamagnetic field against VPMs. In addition, it is expected that the torque density be increased by the additional CPMs. When the magnetizing current is applied, magnetic fluxes pass two main magnetic paths as shown in Fig. 6(c). Hence, the VFMM having double layer CPMs structure may not achieve re-magnetization of VPMs sufficiently because enough magnetizing fluxes aren\u2019t concentrated to VPMs. Therefore, in this paper, extension of the flux barrier is proposed in order to reduce magnetic fluxes which pass the magnetic path including CPMs. The flux barriers that are located nearby VPMs are extended to inner side of the rotor as illustrated in Fig. 6(c). The extended flux barriers can reduce magnetic fluxes which pass the magnetic path including CPMs, and moreover, magnetizing fluxes can be concentrated to VPMs. Consequently, it is anticipated that magnetization characteristics of the VFMM can be remarkably improved by employing double layer CPMs structure and the extended flux barriers. In addition, the extended flux barriers are effective in reduction of the weight of the rotor. IV. COMPARISON OF MAGNETIZATION CHARACTERISTIC In this section, magnetization characteristics of three VFMM models shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002091_0954405420978039-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002091_0954405420978039-Figure4-1.png",
        "caption": "Figure 4. Motions on the machine tool with numerical control for hypoid gear manufacture.",
        "texts": [
            " The loaded tooth contact analysis40 and the thermal mixed elastohydrodynamic lubrication analysis41 are applied to calculate their discrete values. The LTCA and EHL calculations must be run repeatedly through the iteration cycle. The algorithm of the corresponding computer program is shown in Figure 3. The method is fully described in Simon.43 Manufacture of hypoid gears on numerical controlled machine tool The numerical controlled machine tool for hypoid gear manufacture with six degree freedom is shown in Figure 4. There are three rotational motions (u, z, h), and three translational motions (X, Y, Z). The rotational motion, u, around axis xT1, provides the cutting motion of the head-cutter. The other motions are as follows: z is the swinging base rotation angle, h the pinion\u2019s rotation angle, X is the horizontal setting of the cutting spindle, Y is the sliding base setting, and Z is the vertical setting of the cutting spindle. The six axes of the CNC hypoid generator are directly driven by the servo motors and able to implement prescribed functions of motions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002966_14644193211020247-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002966_14644193211020247-Figure1-1.png",
        "caption": "Figure 1. Kinetic model of diesel engine shaft system.",
        "texts": [
            " Alternatively, to use ADAMS/AutoFlex module to generate the flexible bodies, it introduced the meshing capabilities, can generate MNF without the external finite element analysis software, can mesh the geometry generated by external CAD software or ADAMS/View, establish the flexible body. Assembly model of crankshaft system According to a certain type of heavy-duty vehicle diesel engine shafting parts and assembly drawings, establish the crankshaft system consisted of crankshaft, piston, connection rod, crank pin, flywheel, silicone oil damper and other model part, with the three-dimensional modeling software UG, then assemble them shown in Figure 1. Due to the mass and the moment of inertia of the flywheel is very large, relative to the dynamometer and the diesel engine. Flexible connection is adopted in the test of diesel engine bench, the influence of torsional vibration on the crankshaft system is limited. So in the simulation of dynamic response analysis, the dynamometer and the elastic coupling parts is not considered in the rigid flexible mixed multi-body system. Based on the above model, the dynamic response of the system is solved, the time history of the dynamic response of the system is obtained, the torsional vibration response of the system is calculated and analyzed",
            " According to the relationship between the various parts of the shaft system, using ADAMS to establish the corresponding constraint model. Piston pin and piston, piston pin and connecting rod, connecting rod and crank pin, crankshaft and the earth are simplified as hinge constraints. Piston and cylinder suites are simplified as sliding constraints, the belt pulley and the flywheel are respectively fixed with the crankshaft, the shell of silicone oil damper is fixed with the crankshaft, the inertia block and the shell is connected as hinge constraint, add the torsion spring connection as shown in Figure 1. The silicone oil damper Modeling: to establish a fixed connection between shell and crankshaft, to add a rotated connection between inertia block and shell, to add a torsion spring connection between inertia block and shell. Multi-body kinetic model of crankshaft system Using the ADAMS/Auto Flex module transformed the crankshaft into a flexible body. Specific steps are as follows: 1. Create GridUnit type: Tetrahedral units. Unit dimensions: 10mm. 2. Configuring the connection point Use \u201cFind Attachments\u201d option to automatically find the location of connection point between the crankshaft and other parts, select the nodes near the connection points by the cylindrical surface options"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001865_s00170-020-05992-6-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001865_s00170-020-05992-6-Figure3-1.png",
        "caption": "Fig. 3 Punch geometry showing a straight drilled cooling channels and b conformal cooling channels",
        "texts": [
            " The details of the steps involved in the design of conformally cooled tools are beyond the scope of this paper. The punch considered for the study is shown in Fig. 2. The overall tool length required is 745 mm when used in the hot stamping process. The manufacturing and design processes limit the use of the full tool as a single part. Thus, the tool is composed of four inserts with lengths of 225, 150, 150 and 220 mm respectively. Hot stamping tools are usually segmented to allow easier drilling of the conventional cooling channels. Figure 3a shows the punch with straight drilled channels and Fig. 3b shows the one with conformal cooling channels. The model in Fig. 3b shows that the complex sections of the channels are occupying the top section of the inserts. According to the computer-aided design (CAD) model, the height occupied by the complex section is 15 mm. However, previous studies have shown that increasing build heights in the SLM process leads to increased residual stresses [17]. According to Mugwagwa et al. [18], residual stresses can lead to distortions, warping and cracking. Thus, there was a possibility of high stresses if the whole 15 mm height was to be built additively considering the broad width of the inserts (174 mm). Further analysis was done on the model to seek portions that could be machined. From the analysis, it was concluded that only 2.5 mm of the complex structure could be machined. This left a height of 12.5 mm which needed to be built with LPBF. The next stage was to divide the model into two sections: the section to be machined (base body) and that to be built as shown in Fig. 3b. The conformal cooling system design was simulated to compare the performance of the two cooling systems layouts in Fig. 3. PAM-STAMP 2019 was used for the analysis because it can evaluate the cooling system design through temperature calculations [19]. Table 1 shows the simulation parameters used which depict real-life industrial operating conditions. Based on the parameters in Table 1, the simulation is composed of two stages, namely, stamping and cooling. During the stamping stage, the punch encounters the blank and die to perform the deformation. The blank is then held between the punch and die at different times of 2\u20138 s",
            " The temperature map of the punch shows that the conformal cooling tool had a more uniform temperature distribution when compared with the conventional tool as shown in Fig. 5. The tool manufacturing process chain encompassed five stages as summarized in Fig. 6. After the design and simulation step, the CAD geometry for the base body section was used to develop the CAM program for machining the base parts. The next step was to drill the straight part of the cooling channels using a 5 mm TiN-coated carbide drill bit on a 3-axis DMG Mori DMU 65 monoblock milling machine. As seen in Fig. 3, one side of the cooling channels was angled to conform to the shape of the part. Thus, a Hermle C40U 5-axis milling machine was used to drill the angled holes. This process could also have been performed on a 3-axis machine with an angled vice. Figure 7 shows the process of building the top conformal cooling section of the inserts using the laser powder bed fusion process. The machining was done with an overdraft of 1 mm right round and up to the parting line height. This was performed to allow for final machining after the build process and to give room for any distortion that would occur"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002352_j.mechmachtheory.2020.103945-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002352_j.mechmachtheory.2020.103945-Figure7-1.png",
        "caption": "Fig. 7. Two \u201c{4R} [3R] (6R)\u201d deployable units and two \u201c{6R} [3R] (4R)\u201d deployable units.",
        "texts": [
            " Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx Ru1 MP1 MP2 Ru2 Ru4 Ru3 Ru1 MP1 MP2 Ru2 Ru4 Ru3 Ru1 MP1 MP2 Ru2 Ru4 Ru3 Ru2 Ru1 Ru3 Ru4 MP1 MP2 (a) (b) (c) (d) W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx 5 Further, some units should be excluded due to their structural defects, such as local DOFs and undesired angle ranges[35]. Finally, there are fifteen available \u201c{4R} [3R] (6R)\u201d deployable units, listed in Table 2 , in which two typical units are shown in Fig. 7 (a) and (b). Similarly, seventeen available \u201c{6R} [3R] (4R)\u201ddeployable units can be obtained, listed in Table 3 , in which two typical units are shown in Fig. 7 (c) and (d). Based on the combination of three kinds of \u201c{6R}\u201d modules, one kind of \u201c[2R]\u201d module and three kinds of \u201c(5R)\u201d modules, there are nine \u201c{6R} [3R] (5R)\u201d units in total. After excluding those units with structural defects, six available\u201c{6R} [2R] (5R)\u201d units can be obtained, listed in Table 4 , in which two typical units are shown in Fig. 8 (a) and (b). Similarly, there are also six available \u201c{6R} [1R] (6R)\u201ddeployable units, listed in Table 5 , in which two typical units are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002650_s12555-019-1004-6-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002650_s12555-019-1004-6-Figure1-1.png",
        "caption": "Fig. 1. Four-wheeled omni robot.",
        "texts": [
            " Then the system\u2019s overall stability is proven based on Lyapunov method in Section 3. To verify the accuracy and effectiveness of the proposed controller, Section 4 presents some simulation scenarios. Moreover, this section also demonstrates numerous comparisons between the results of the proposed control algorithms to show the superior of the proposed control strategy. Finally, conclusions are given in Section 5. 2. MATHEMATICAL MODEL OF OMNI ROBOT The Omni-directional mobile robot used for developing control design consists of four Omni wheels that are 90\u25e6 apart as in Fig. 1. Oxy coordinate is chosen as the global coordinate frame. The distance between wheels and robot centre is defined as d. Meanwhile, the robot\u2019s velocity comprises vx, vy and \u03c9 , respectively. State variables are the position of the robot in the global coordinate and the robot\u2019s velocites. 2.1. Kinematic model To determine the robot\u2019s movement, a position vector is defined as q = [ x y \u03b8 ]T and a velocity vector on the fixed frame is the derivative of q. Through the following equation, the mentioned velocity can be turned into the velocity on the robot\u2019s frame q\u0307 = Hv, (1) where v = [ vx vy \u03c9 ]T denotes the velocities of the robot and the transition matrix H = cos\u03b8 \u2212sin\u03b8 0 sin\u03b8 cos\u03b8 0 0 0 1  demonstrates the relationship between q and v"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000567_j.jsv.2016.04.020-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000567_j.jsv.2016.04.020-Figure7-1.png",
        "caption": "Fig. 7. Axial subsystem and its four sub-subsystems.",
        "texts": [
            " Connecting the front bearing block C via rotational damping, yields: HABC\u00f0l\u00de \u00bc h0r00 h0r0l\u00f0l\u00de h0r0L h0rl0\u00f0l\u00de h0rll\u00f0l\u00de h0rlL\u00f0l\u00de h0rL0 h0rLl\u00f0l\u00de h0rLL 2 64 3 75 \u00bcHAB\u00f0l\u00de hr00 hrl0\u00f0l\u00de hrL0 2 64 3 75 hr00\u00fe0\u00fe 1 ic\u03b8b1\u03c9 1 hr00 hr0l\u00f0l\u00de hr0L 2 64 3 75 T (27) Connecting the rear bearing block D via rotational damping, yields: HABCD\u00f0l\u00de \u00bc h\u2033r00 h\u2033r0l\u00f0l\u00de h\u2033r0L h\u2033rl0\u00f0l\u00de h\u2033rll\u00f0l\u00de h\u2033rlL\u00f0l\u00de h\u2033rL0 h\u2033 rLl\u00f0l\u00de h\u2033rLL 2 664 3 775 \u00bcHABC\u00f0l\u00de h0 r0L h0 rlL\u00f0l\u00de h0rLL 2 64 3 75 h0rLL\u00fe0\u00fe 1 ic\u03b8b2\u03c9 1 h0rL0 h0rLl\u00f0l\u00de h0rLL 2 64 3 75 T (28) Lastly, connecting the motor rotor E via the coupling, yields: HRot\u00f0l\u00de \u00bcHABCDE\u00f0l\u00de \u00bc hRmm hRml\u00f0l\u00de hRlm\u00f0l\u00de hRll\u00f0l\u00de \" # \u00bc hm 0 0 h\u2033rll\u00f0l\u00de \" # hm h\u2033rl0\u00f0l\u00de \" # h\u2033r00\u00fehm\u00fe 1 kc\u00fe i\u03c9cc 1 hm h\u2033r0l\u00f0l\u00de \" #T (29) where, the subscript \"R\" means rotational, \u201cm\u201d means the coordinate at the motor rotor, and \u201cl\u201d the coordinate at the nut location on the ballscrew. It is worth noting that explicit function with respect to the nut location l is maintained in Eq. (29). The schematic model of the axial subsystem is given as the top half of Fig. 7. The ballscrew is axially supported with the front and rear bearings, which apply axial stiffness and damping kb1, cb1, kb2, cb2 on the ballscrew's ends. Similar to the rotational subsystem, axial subsystem is divided into four sub-subsystems A, B, C, D as illustrated in the bottom half of Fig. 7. Using a very similar approach as in Section 3.1.1, the direct FRF at the ballscrew's coordinate l can be obtained, as Eq. (30). HAxl\u00f0l\u00de \u00bcHABCD\u00f0l\u00de \u00bc hAll\u00f0l\u00de (30) where, the subscript \"A\" means axial, \"l\" means the ballscrew's coordinate at the nut location. The structure of the sliding subsystem is given in Fig. 8, where, A is the carriage assumed rigid supported on Y guide and driven by a ballscrew; B is the table assumed rigid supported on X guide that mounted on the carriage; 1, 2, 3 is the coordinates located at the center of nut, the scan head of the linear scale, and the center of the edge of the table; h1, h2, h3, and h4 are vertical distances between the coordinates, the centers of mass and the center plane of the X guideblocks"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003709_s0043-1648(96)07486-8-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003709_s0043-1648(96)07486-8-Figure2-1.png",
        "caption": "Fig. 2. Outline sketch of a spherical roller thrust bearing.Fig. 1. Slip lines in an elliptical contact.",
        "texts": [
            " The arrows shown in the contact represent the directions of slip when the ball is rolling into the paper. There are two points along the contact where the sliding velocity is zero. At all other points along the contact, sliding is present. Zero-sliding points have been observed in spherical roller bearingsbyFysh et al. [3]. In an experimental study Brothers and Halling [4] found that the conformity ratio affects the amount of wear in a conform rolling sliding contact. Spherical roller thrust bearings (see Fig. 2) consist of four parts: the shaftwasher, the housingwasher, the cageassembly and the rollers. These bearings can accommodate heavy combined loads and are insensitive to misalignment of the shaft relative to the housing. Spherical roller thrust bearings are used in machinery such as gear boxes, wind power stations, electric motors, pumps, hydraulic motors and cranes. More information about spherical roller thrust bearings can be found in [5]. Under boundary lubricated conditions these bearings can break down due to rupture at the two circular bands around the washer surfaces (see Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002764_0142331221994393-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002764_0142331221994393-Figure4-1.png",
        "caption": "Figure 4. Turning diagram of the car-like model.",
        "texts": [
            " Time optimization is another constraint considered by TEB, which requires the time interval among all pose positions to be minimized in order to reach the target point within the shortest time and subsequently improve efficiency. The time optimal constraint function is expressed as given by Equation (13): f =( Pn i= 0 DTi) 2 i 2 N \u00f013\u00de The real-time motion state of the car-like robot is represented by Equation (14), where b(t) is the steering angle of the rear axle centre. The TEB approach is used for local path planning. The trajectory between each pose is approximately regarded as an arc, and the combination of multiple trajectories is the local optimal path planned by TEB. Figure 4 shows a car-like robot model with two consecutive poses, where Dbi is the real-time steering angle of the rear axle centre. The car-like robot travels on the planned path when turning. Two consecutive adjacent motion poses are located in a common arc with constant curvature that satisfies ui, i = ui, i+ 1 and di = xi+ 1 xi yk + 1 yk 0\u00bd T . By connecting the two poses, the non-holonomic kinematic constraint can be formulated as given by Equation (15): _S t\u00f0 \u00de= _x t\u00f0 \u00de _y t\u00f0 \u00de _b t\u00f0 \u00de 2 4 3 5; _b t\u00f0 \u00de= v L tan aa t\u00f0 \u00de\u00f0 \u00de \u00f014\u00de fi(si+ 1, si)= cos (bi) sin (bi) 0 2 4 3 5+ cos (bi+ 1) sin (bi+ 1) 0 2 4 3 5 0 @ 1 A3 di = 0 \u00f015\u00de The minimum turning radius and maximum turning angle are related to the structure of the car-like robot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000998_tmag.2019.2902428-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000998_tmag.2019.2902428-Figure7-1.png",
        "caption": "Fig. 7. Distributions of hysteresis loss density. (a) 1-D FEM. (b) Methods A, B, and C. (c) Method D.",
        "texts": [],
        "surrounding_texts": [
            "We investigated the Cauer circuit modeling methods of dynamic hysteretic property in iron loss analyses of practical electric machines as the post-processing of the main magnetic field analysis. As a consequence, the computational accuracy of the Cauer circuit models is almost the same as the ordinary 1-D FEM including the ratio of eddy-current loss to hysteresis loss, although the computation time of Cauer circuit models is much lower. Therefore, the Cauer circuit models are one of the hopeful options to estimate the iron loss of electric machines within the acceptable computational cost."
        ]
    },
    {
        "image_filename": "designv11_14_0001834_j.addma.2020.101547-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001834_j.addma.2020.101547-Figure4-1.png",
        "caption": "Fig. 4. Schematic of the laser path (a) first layer, (b) second layer, (c) working position of the laser on the substrate.",
        "texts": [
            " In this study, a diode laser with a wavelength of 1064 nm was used as the energy source and argon shielding gas was applied to protect the surfaces of product in the DED process. The laser type is LDF 3000-60 VGP, which is provided by Laserline GmbH. The major process parameters were listed in Table 1. Inconel 625 was chosen as the target material. The chemical composition and particle size of the Inconel 625 are listed in Table 2 and Table 3, respectively. The AM part used in the experiment was a disc-shaped specimen. The deposition track of the disc-shaped specimen (2 r= 25.0mm) in the DED process is shown in Fig. 4: hatch spacing was w=2.0mm and there were 11 fusion lines in each layer. Figs. 4a and 4b show the deposition tracks of the first and second layers, respectively. As shown in Fig. 4c, the deposition orientations of the first and second layers were perpendicular to each other. The substrate used in this study is a thick substrate with a thickness of 16mm, which is made of 304-stainless steel. As shown in Fig. 5, the bottom of the substrate was polished. As shown in Fig. 6, a special fixture is designed in the experiment. The substrate was coaxially clamped on the center hole of the support frame by two bolts. The fixture consists of bolts, support frame and springs. The gap between the substrate and the fixture ensures that the substrate can move freely in the x- and y-directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002542_icem49940.2020.9270965-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002542_icem49940.2020.9270965-Figure6-1.png",
        "caption": "Fig. 6. Interspersed winding and prototype [22].",
        "texts": [
            " The Odd phase numbers have higher torque ripple frequency and lower amplitude characteristics, while the even phase number may be beneficial in compensating for unbalanced magnetic pull when a phase is disabled. In the practical application of MEA, the optimal phase number may differ due to the varying subsystem requirements and limitations. An electric drive system based two sets of independent three-phase concentrated armature windings FT PM motor for aerospace applications is proposed and investigated [21], with the merits of high power density, high efficiency and high fault tolerance. Dual three-phase PM motor for safety critical aircraft actuator has been designed and tested [22] (Fig. 6.) which has a higher slot fill factor, better manufacturability by rectangular winding conductor and lower probability of turnturn faults. To avoid the reliability over-designed and give consideration to technology readiness level, dual three-phase topology is the common one in electrical machine phase\u2019s selection. The feasibility and fault tolerance of slot/pole combinations are evaluated in the following sections. D. Innovative Topologies Some innovative topologies have been presented based on modular design, flux controllable and separated phase topologies"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000719_mma.4210-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000719_mma.4210-Figure1-1.png",
        "caption": "Figure 1. Rotating extensible robotic arm system model.",
        "texts": [
            " In this paper, a mathematical approach for generating the smooth trajectory of the end effector of a rotating extensible robotic arm is presented. Cartesian and polar piecewise polynomial interpolating curves are considered for the generation of the geometric path of the end effector. To verify the proposed approach, the trajectory and the velocity profile for the end effector and non-active end of the constrained trajectory are computed for two different configurations. 2. Mathematical modelling of manipulator trajectory 2.1. System model The rotating extensible robotic arm shown in Figure 1 is composed of a rigid guide OE and a sliding rod SP, which is constrained to a curved trajectory by the end S. The non-active end of the sliding part SP is denoted by S, and its active end-effector is denoted by P. The rigid guide of the robotic arm has length dOE ; the sliding part has the length dSP . The total length of the robotic arm denoted by r D dOP (distance between the manipulator base location O and its end-effector P/ varies because of the rotation of the rigid guide of Department of Design and Engineering, Faculty of Science and Technology, Bournemouth University, Poole, UK * Correspondence to: Mihai Dupac, Department of Design and Engineering, Faculty of Science and Technology, Bournemouth University, Poole, UK"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001999_ecce44975.2020.9235600-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001999_ecce44975.2020.9235600-Figure1-1.png",
        "caption": "Fig. 1. Cross-sectional views of two different VFMMs.",
        "texts": [
            " In many conventional VFMMs, larger positive magnetizing current pulse are required than negative one because re-magnetization is difficult to concentrate magnetic flux to VPMs due to the saturation and the diamagnetic field from constant flux PMs (CPMs). Eventually, asymmetric positive and negative magnetizing current causes increase in the inverter capacity. In addition, unintentional demagnetization in VPMs tends to be caused by load current. Unintentional demagnetization is not desirable in terms of reliability. Furthermore, demagnetized VPMs might cause large iron loss due to the harmonic flux. In literature [4] which is reported by our research group, the VFMM employing delta-type PM arrangement and extended flux barriers shown in Fig. 1(a) can overcome above problems. At the same time, the VFMM proposed in [4] indicates higher efficiency than that of an interior permanent magnet synchronous motor (IPMSM) mounted in TOYOTA PRIUS 4th generation that is commercially supplied HEV in 2015 [17]. However, this conventional VFMM needs 200 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 53 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 08:13:29 UTC from IEEE Xplore. Restrictions apply. Arms to achieve target maximum torque of 163 Nm because this model cannot generate reluctance torque effectively",
            " Large maximum current causes large copper loss and needs thick wire harness. Accordingly, compared with the target IPMSM, efficiency around maximum torque region is lower and total weight of wire harness might be increased, especially traction applications. In this paper, in order to resolve above problems of the conventional VFMM, improved VFMM which can effectively generate the reluctance torque to achieve target value of 163 Nm by same maximum load current as target IPMSM is proposed as shown in Fig. 1(b). Another important purpose in this paper is to reduce the maximum magnetizing current pulse. 2D-FEA results shows that the proposed VFMM can achieve target torque of 163 Nm at 180 Arms and reduce maximum magnetizing current pulse by 60 Arms compared to the conventional VFMM. II. CONCEPT OF PROPOSED AND CONVENTIONAL VFMMS Table I lists target specifications of the conventional VFMM reported in [4] and the proposed VFMM. Most of the parameters of the conventional and proposed VFMMs are same as the target IPMSM mounted in TOYOTA PRIUS 4th generation",
            " The conventional VFMM needs 200 Arms, which is higher than that of target IPMSM, to achieve the maximum torque of 163 Nm because the magnetic torque is dominant. In large current application like vehicle traction, to increase the reluctance torque is very effective in improving torque density since the reluctance torque is proportional to the square of load current. Accordingly, in this paper, a useful design of VFMM which can use the reluctance torque effectively is proposed in order to achieve the maximum torque of 163 Nm by same load current as the target IPMSM. Fig. 1 shows crass-sectional views of conventional and proposed VFMMs. Both models employ delta-type PM arrangement and extended flux barriers in order to improve magnetization characteristics [4]. In addition, they have same stator shape as the target IPMSM that uses the segmented rectangular winding. Moreover, VPMs, which have variable magnetization ratio, are installed between poles in the rotor of both VFMMs. In this case, commercially-supplied Sm-Co material is used for VPMs. In addition, Neodymium sintered PM is employed for CPMs that configurate delta-type PM arrangement"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001531_1.c035829-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001531_1.c035829-Figure2-1.png",
        "caption": "Fig. 2 Tire force analysis diagram.",
        "texts": [
            " They are calculated in Sa \u2212Oaxayaza: Lt 0 0 \u2212 1 2 \u03c1V2 t StCL\u03b1;t\u03b7e\u03b4e T (8) Pr 0 1 2 \u03c1V2 rSrC\u03b4r\u03b7r\u03b4r 0 T (9) Fc \u2212F 0 0 T Fc 8>>>>>>< >>>>>>: 1 2 \u03c1V2ScCLct1 0 \u2264 t1 \u2264 1 1 2 \u03c1V2ScCLc t1 > 1 & V > vd 0 V \u2264 vd (10) The corresponding moments can be expressed in Sb \u2212Obxbybzb: Mc 0 Fccd \u2212 Ltct \u2212Prlr T (11) where Vt and Vr are the velocities at the horizontal and vertical tails; St, Sr, and Sc are the elevator, rudder, and drag parachute areas; \u03b7e and \u03b7r are the efficiencies of the elevator and the rudder control surfaces; \u03b4e and \u03b4r are the control surface deflections; t1 is the rollout time from the moment the aircraft touches the ground; cd, ct, and lr are the perpendicular distances between the force points on the three surfaces and the UAV gravity center; and CL\u03b1:t, C\u03b4r , and CLc are the lift, side force, and drag coefficients of the elevator, rudder, and drag parachute respectively. The relationships between them and \u03b1 and \u03b2 are similar to those in Figs. 1a and 1c. The tire lateral force is a nonlinear function of the tire slip angle and ground reaction force. The tire slip angle is defined as the inclined angle of the tire symmetry plane and the tire traveling direction. The force analysis diagramsof the nose tire and themain tire are illustrated in Fig. 2. \u03b8l is the nose wheel steering angle controlled by a steering actuator. The slip angles of the nose wheel and the two main wheels \u03b2n; \u03b2ml; \u03b2mr can be calculated as \u03b2n 8>>>>>>>>< >>>>>>>>: arctan vsy r \u22c5 an vsx \u2212 \u03b8l vxn \u2265 0 \u03c0 arctan vsy r \u22c5 an vsx \u2212 \u03b8l vxn < 0 & vyn \u2265 0 \u2212\u03c0 arctan vsy r \u22c5 an vsx \u2212 \u03b8l vxn < 0 & vyn < 0 (12) \u03b2ml 8>>>>>>>>< >>>>>>>>: arctan vsy \u2212 r \u22c5 am vsx r \u22c5 bw\u22152 vxml \u2265 0 \u03c0 arctan vsy \u2212 r \u22c5 am vsx r \u22c5 bw\u22152 vxml < 0 & vyml \u2265 0 \u2212\u03c0 arctan vsy \u2212 r \u22c5 am vsx r \u22c5 bw\u22152 vxml < 0 & vyml < 0 (13) \u03b2mr 8>>>>>>>>< >>>>>>>>: arctan vsy \u2212 r \u22c5 am vsx \u2212 r \u22c5 bw\u22152 vxmr \u2265 0 \u03c0 arctan vsy \u2212 r \u22c5 am vsx \u2212 r \u22c5 bw\u22152 vxmr < 0 & vymr \u2265 0 \u2212\u03c0 arctan vsy \u2212 r \u22c5 am vsx \u2212 r \u22c5 bw\u22152 vxmr < 0 & vymr < 0 (14) where vsx, vsy, and vsz are the UAV velocity components in the stability axis system Ss \u2212OsXsYsZs; r is the yaw rate; an and am are the distances form the UAV gravity center to the nose gear and the main gears, respectively; and bw is the main-wheel span"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001217_j.mechmachtheory.2019.103633-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001217_j.mechmachtheory.2019.103633-Figure8-1.png",
        "caption": "Fig. 8. Deformation analysis of 4BSL under force F \u03b3 .",
        "texts": [
            " To overcome the disadvantages of the spring-crank mechanism, a negative stiffness mechanism\u2014spring four-bar linkage (4BSL) is designed in this paper, as shown in Fig. 5 . This mechanism is a diamond-shaped four-bar linkage, which is composed of four equal-length linkages and a spring at the diagonal position. The 4BSL is applied between the inner and outer ring of the IORFH. The inner ring is equivalent to a crank. So, a crank spring four-bar linkage mechanism (Crank4BSL) is formed, as shown in Fig. 6 . Consequently, a kind of new ZSFH based on 4BSL (ZSFH_4BSL) is designed, as shown in Fig. 7 . As shown in Fig. 8 , the 4BSL is deformed from B \u03b2 to B \u03b3 by the force F \u03b3 , and the amount of deformation is x \u03b3 . \u03d5 is the angle between B \u03b3 Q \u03b3 and x -axis, e is the length of the linkage, F ls is the restoring force of the spring, and F link is the internal force of the linkage. By the force received analysis of B \u03b3 and Q \u03b3 , we can get: F \u03b3 = cot \u03d5 F ls = cot \u03d5 K 4 BSL _ S x ls (1) Where: K 4BSL_S is the stiffness of the spring of 4BSL, x ls is the deformation of the spring. From the geometric relation- ship, we can know: \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 x \u03b3 = \u2223\u2223B \u03b2C \u2223\u2223 \u2212 \u2223\u2223B \u03b3 C \u2223\u2223 x ls = \u2223\u2223Q \u03b3 P \u03b3 \u2223\u2223 \u2212 \u2223\u2223Q \u03b2P \u03b2 \u2223\u2223\u2223\u2223Q \u03b2P \u03b2 \u2223\u2223 = 2 \u221a e 2 \u2212 ( 0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002876_tmech.2021.3072939-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002876_tmech.2021.3072939-Figure5-1.png",
        "caption": "Fig. 5. (a) Uni-directional testbed to analyze string parameters and measure NSTR,PBDA, and NFTR. (b) A single motor driven bidirectional robotic joint utilizing PBDA and ABC mechanisms.",
        "texts": [
            " 16, as follows: hABC \u2212 rtendon(cos\u03c8 + 2hABC sin\u03c8 xABC ) < rmin (17) If we select both controllable variables hABC and xABC to satisfy the above inequality, Eq. 17, we can prevent overtorsion. A larger xABC and a smaller hABC are preferable for over-torsion prevention. xABC selects the upper limit using the size of the ABC and \u03bamax. The overall design flowchart of the ABC is shown in Fig. 4. For the experimental verification of the proposed method, we developed a testbed to evaluate the performance of the selected design parameters (see Fig. 5(a) and Table I). We used a Maxon 3.9:1 DCX32 motor (102 W) to twist the strings Authorized licensed use limited to: Konya Teknik Universitesi. Downloaded on May 18,2021 at 14:15:55 UTC from IEEE Xplore. Restrictions apply. 1083-4435 (c) 2021 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. and installed a Maxon 74:1 RE40 motor (150 W) at the joint to apply the external load. The US Digital E3-10000 CPR rotary encoder was placed at the joint to measure the angular displacement. The UNIPULSE UTM2-50 was installed at the joint to measure the output torque, as shown in Fig. 5(a). The US Digital EM2-2000 linear sensor was placed at the end of strings to measure linear displacement. The CASKOREA CBF30-200 was placed at the ends of the strings to measure the string forces, as shown in Fig. 5(b). The end effector was 0.2 m long and weighed 0.2 kg. For controlling the actuators, position feedback control and feedforward control are implemented, as shown in Fig. 6. For real-time application, we adopted a MAXPOS EtherCAT motor driver with a STEP-PC2 (Neuromeka) real-time operational system PC at a sampling rate of 1 KHz. We conducted the system identification process to identify the complex dynamics of TSAs. We used sinusoidal current input at various frequencies and measured the output torque",
            " In order to evaluate the dynamic performance of both TSAs, we conducted the position tracking with full RoM (\u00b190\u25e6) and sinusoidal command (\u00b15\u25e6, 4.0 Hz). For conventional TSA, position root-mean-square errors (RMSE) were measured to be 0.7\u25e6 for RoM experiment and 0.4\u25e6 for control bandwidth experiment at 4.0 Hz. For proposed TSA, RMSE were measured to be 0.6\u25e6 (RoM) and 0.9\u25e6 (control bandwidth), respectively. B. Validation of Parallel Bundle\u2013Driven Actuation First, we investigate a characteristic of the NFTR employing the testbed shown in Fig. 5(a). Then, we compare the input energy consumption of TSAs (same L) under various output conditions employing the weights as shown in Fig. 5(b). Fig. 8 shows that NSTR increases as Nbundle increases. The experimental results showed close resemblance to the simulation results. B3 exhibited the highest NFTR, which is 21% greater than that of B1. B6, with the smallest R of the four cases, showed the highest NSTR. However, B6 exhibited the lowest \u03b7PBDA (50.7%) because of the energy loss by additional transmission such as gears and strings. The selected B3, which has the highest NFTR, showed that NSTR was higher than B1 in all TSA configurations, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002072_j.simpat.2020.102236-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002072_j.simpat.2020.102236-Figure2-1.png",
        "caption": "Fig. 2. Schematic view of misaligned journal bearings.",
        "texts": [
            " \u2202 \u2202t \u222b \u03a9(t) \u03c1fd\u03a9 + \u222b \u03c3 \u03c1 ( ( up \u2212 ug ) \u22c5 n ) fd\u03c3 = \u222b \u03c3 ( Df + \u03bct \u03c3f ) (\u2207f \u22c5 n)d\u03c3 + \u222b \u03a9 (Re \u2212 Rc)d\u03a9 (3) Re = Ye \u0305\u0305\u0305\u0305 K \u221a \u03c3\u03b9 \u03c1\u03b9\u03c1\u03bd [ 2 3 (P \u2212 P\u03bd) \u03c1\u03b9 ]1/2( 1 \u2212 fv \u2212 fg ) Rc = Yc \u0305\u0305 k \u221a \u03c3\u03b9 \u03c1\u03b9\u03c1\u03bd [ 2 3 (P \u2212 P\u03bd) \u03c1\u03b9 ]1/2 fv (4) The mixture density \u03c1mix of the fluid can be obtained by: 1 \u03c1mix = fv \u03c1v + fg \u03c1g + 1 \u2212 fv \u2212 fg \u03c1\u03b9 (5) The turbulence in the pump will influence the pump performance at a high Reynolds number state. For a long computational time simulation, the higher order turbulence model will cost more simulation time without an obvious improvement in the results [45]. Thus, the standard k-\u03b5 model [1,46] is used in the pump CFD model because of its satisfactory robustness, efficiency and accuracy. The numerical model of EGP is built using the commercial code PumpLinx. Fig. 2 is the schematic view of the misaligned journal bearings. Using three parameters (\u03b1,e1and Dm) to define the misalignment, the oil film thickness can be presented as follows [25]: Note that the misalignment considered is supposed to be constant both in magnitude and direction. J. Zhu et al. Simulation Modelling Practice and Theory 108 (2021) 102236 h = hgeo + e1(z /L \u2212 0.5)cos(\u03b8 \u2212 \u03b1 \u2212 \u03d50) (6) The geometric height hgeo can be calculated as: hgeo = C + ecos(\u03b8 \u2212 \u03d50) (7) The emax is the maximum possible value of e1, which can be obtained by: emax = 2 [( C2 \u2212 e2 0sin2\u03b1 )0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001535_j.ymssp.2020.106723-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001535_j.ymssp.2020.106723-Figure4-1.png",
        "caption": "Fig. 4. Schematic diagram of motion modes and acceleration components.",
        "texts": [
            " The existence of offsets h2 and l2 forced line AB to rotate around line AO with a certain angle g. The angle is related to distance of the offset and coordinate of point E. The expression of angle g is: g \u00bc arccos l2 h2\u00f0 \u00de= ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi X2 E \u00fe YEsin c\u00f0 \u00de ZEcos c\u00f0 \u00de\u00f0 \u00de2 q p=2 \u00f01\u00de A relative rectangular coordinate at G is constructed as shown in Fig. 4. Defining unit vector of l ! AB as e!AB, the choice of origin of the relative coordinate is set at a random point on line OE. Points G and F are intersections of lines AD and OE in plane Q , where plane Q is parallel with plane OAB and crosses at point G. In relative coordinate system, x-axis is parallel with X-axis in absolute coordinate system, y-axis is along with line GF, and z-axis is perpendicular to x-axis and y-axis. The unit vectors of x-axis, y-axis, and z-axis are i\u2019, j\u2019, and k\u2019, respectively. When the cardan shaft rotates with a certain misalignment, the rotating motion contains two parts, one is the relative motion and the other is transport motion. Assuming point Q is the intersection of plane Q and line BC, the kinematic model relating to point Q is shown in Fig. 4. The motion of point Q also contains two motions, one is the transport motion rotating around the x-axis and the other is the relative motion rotating around the y-axis. With an angular speed of drive shaftx; the angle between line AB and plane YOZ is /, and the velocity of point Q is expressed as: vQ \u00bc v transport \u00fe v relative \u00bc xQ di0 dt \u00fe yQ dj0 dt \u00fe zQ dk0 dt \u00fe dxQ dt i0 \u00fe dyQ dt j0 \u00fe dzQ dt k0 \u00bc x! l ! GQ \u00fe d/=dt\u00f0 \u00de e!GF l ! FQ \u00f02\u00de The acceleration of point Q contains 4 parts, namely, normal transport acceleration, tangential relative acceleration, normal relative acceleration and Coriolis acceleration"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001934_iros45743.2020.9341409-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001934_iros45743.2020.9341409-Figure9-1.png",
        "caption": "Fig. 9: Simulation results for planning with tether constraints. In test 1 we demonstrated the planner\u2019s ability to plan safe paths when there is a stable anchor. In test 2 we moved the anchor to an unstable position and demonstrated the planner\u2019s ability to recognize the slope as hazardous. In test 3 we added a pole and demonstrated the planner\u2019s ability to find intermediate, stable anchor points.",
        "texts": [
            " For each test, after 750 iterations the planner was unable to find a safe path to the goal. These two tests in conjunction with each other demonstrate the importance of anchor point placement on traversability. In the final test we add a pole to the middle of the map. In this configuration we expect the planner to find a path that wraps around and anchors to the pole before descending. Out of the 100 trials performed, 84 of the tests found paths in the stable homotopy class of anchoring to the pole and 16 of the tests resulted in no paths found. Paths for each test can be seen in Fig. 9. These figures show the start and goal positions of the robot, and their respective anchor histories. Green lines indicate found paths. For test1, it shows that the planner recommended the straight line down the slope and an intermediate anchor was found on the lip of the slope as the rover descended. For test2, it shows that ABIT* has correctly failed to find a safe path to the bottom. For test3, it shows that the path has the rover drive around the pole, generating an intermediate anchor point that allows it to drive directly down the slope"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003688_1.1304914-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003688_1.1304914-Figure1-1.png",
        "caption": "Fig. 1 Schematic of the Dynabee. The precessional motion of a vector normal to the circumferential track is also shown.",
        "texts": [
            " In this section, we discuss the kinematics of the proposed mechanical model of the Dynabee. This model contains the important physical characteristics necessary for spin-up to occur. Specifically, bases vectors and Euler angles are introduced to describe the rotations of two bodies: a circular track and a rotor. For the relevant background on parameterizations of rotation tensors, the reader is referred to the review article by Shuster @5#. The mechanical model is comprised of two rigid bodies, a track and a rotor, as illustrated in Fig. 1. The track is a circular race or groove that constrains the motion of the rotor, which is an axisymmetric body with a cylindrical axle along its axis of symmetry. With a semi-length of Rt and a radius of Ra , both ends of the axle are constrained to remain within the track. The rotor\u2019s moment of inertia about a vector parallel to the axle, referred to as the rotor\u2019s spin-axis, is l1 . Due to a geometric symmetry, the moment of inertia about any vector perpendicular to the spin axis is l2 . 1To whom correspondence should be addressed"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001829_s12541-020-00389-7-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001829_s12541-020-00389-7-Figure12-1.png",
        "caption": "Fig. 12 Experimental setup of static stiffness test",
        "texts": [
            " Unlike the numerical model, the 1 3 actual spindle is placed with peripherals, such as a housing and a fixture, which could make the whole structure weaker. Table\u00a05 presents the predicted static stiffness values of the initial and optimal design spindles. The static stiffness values of the optimal spindles are all higher than those of the initial spindle, regardless of the applied preload. Exceeding 26% is obtained when comparing the static stiffness of the optimal spindle and that of the initial spindle. Figure\u00a012 shows the experimental setup for the static stiffness test. In this test, the preload varies from 900 to 2000, and to 3000\u00a0N under non-rotating conditions. An external load in the vertical direction (Y) is applied to the spindle nose using a hydraulic jack, as shown in Fig.\u00a012. The applied radial load is measured by a compression-type load cell and the associated radial displacement is acquired by a lever-type inductive probe at point A, as indicated in Fig.\u00a012. Figure\u00a013 shows a comparison of the measured and predicted stiffness values of the spindle, under a moderate axial preload (2000\u00a0N). The measured and calculated displacements of the spindle match each other well (Fig.\u00a013b\u2013c). In the case of a low preload of 900\u00a0N, the measured and predicted displacements are in good agreement with the radial load, up to approximately 2000\u00a0N. Beyond this load, an apparent difference is observed between the measurement and prediction (dashed line in Fig.\u00a013a). In the case of a light preload, for example, 900\u00a0 N, the preload is generated only by a set of helical springs on the right side, without any additional hydrostatic force from the compressed air"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000005_s10659-019-09744-w-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000005_s10659-019-09744-w-Figure5-1.png",
        "caption": "Fig. 5 Illustrations of Sauer\u2019s approach to define discrete vertex based Frenet frames of a polygonal arc, taken from \u00a72 of his book [54]: Fig. 2.1 explains the definition of vertex based tangent, binormal and principal normal vectors, and vertex angles between tangent vectors; Fig. 2.2 illustrates the definition of the rotation angle of discrete osculating planes, equal to the angle enclosed by the corresponding binormals",
        "texts": [
            "10 The latter fact makes the (unit length) vector (cj \u2212 pj )/\u03c1j an alternative candidate for the definition of a discrete principle normal located at pj , which implies an alternative definition of a discrete vertex tangent vector. These become identical to n\u0304j and t\u0304j for j\u22121/2 = j+1/2, but in general are different otherwise. Our approach to the subject may be considered as a reinterpretation of Sauer\u2019s approach, as presented in Sect. 2 of his book [54], in terms of staggered finite differences, rather than a vertex based view point taken by Sauer (see Fig. 5). As explained below, by our modifications of discrete curvature definitions and the treatment of discrete principle normals we are able to avoid the appearance of extra terms in the discrete Frenet equations that vanish only in the continuum limit. Vertex Based Discrete Frenet Frames While Sauer uses the same definition (i.e.: b(S) j \u2261 bj ) of vertex based discrete binormals as presented here, he interprets the definition of discrete tangents along edges as a (likewise vertex based) forward difference t(S) j \u2261 tj+ 1 2 , and consistently introduces vertex based principal normals as n(S) j := b(S) j \u00d7 t(S) j \u2261 bj \u00d7 tj+ 1 2 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure45-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure45-1.png",
        "caption": "Fig. 45. Legge Synchro-Coupling",
        "texts": [
            " When starting, some oil leakage would take place through the sleeve shaft into the gearbox, hence the latter is provided with a sump pump driven from the output shaft and piped to the oil return connexion from the cooler, so that any excess of oil in the sump would be returned to one of the fluid couplings. An equalizing port is arranged between the reservoirs to permit any excess in one to overflow centrifugally into the other. The couplings have no glands, and when stationary, the oil would be retained in the lower part of the casings. Legge Synchro-Coupling. The principle of the Legge synchrocoupling used to prevent backward rotation of the annulus when running in low gear will be understood from Fig. 45, which illustrates a synchro-coupling arranged to transmit power between two shafts. In this construction the input shaft on the left is permitted to run slower than the output shaft on the right when the synchro-coupling is disengaged, as drawn. On accelerating the input shaft the synchro-coupling engages automatically and positively upon synchronism. A locking ring is shown in this example to give a bi-directional drive when desired. A synchro-coupling consists essentially of the following : (1) An input shaft carrying a ring of dogs A, also a pair of pawls B for picking up similar dogs on an intermediate member when synchronism is reached",
            " These dogs then prevent relative rotation between the intermediate member C and the driven shaft E and thus lock the synchro-coupling to give bi-directional drive. Since the synchro-coupling in Fig. 42 is only used to prevent backward rotation of the annulus, there is no locking sleeve on the intermediate member which, it will be observed, is mounted on a multi-start thread on the boss of the annulus. The stationary dogs in this example correspond to the driving dogs on the input shaft in Fig. 44. When the synchro-coupling is out of engagement as in Fig. 45, the input shaft on the left can rotate in the direction of the arrow, slower than the output shaft, and the pawls are then overrun by the dogs on the helically mounted intermediate member. When the input shaft is accelerated, the pawls attain synchronism with the dogs on the intermediate member, with which they engage quietly, and thus displace it axially into precise synchronous engagement with the driving dogs. When the dogs are in full mesh the intermediate member is prevented from further axial movement by a stop, and the synchro-coupling is then capable of transmitting torque, the pawls being entirely free from load"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001218_s42417-019-00182-5-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001218_s42417-019-00182-5-Figure7-1.png",
        "caption": "Fig. 7 Diagram showing the ball screw temperature distribution",
        "texts": [
            " Among them, the temperature values for Q1\u2013Q5 are obtained by from the heat model in Eq.\u00a0(6). Based on the heat milling thermal theory model, the temperature of the cutting edge reaches a steady state after approximately 0.6\u00a0s during high-speed milling. The results show that the milling temperature at the cutting edge fluctuates around 50\u00a0\u00b0C. Therefore, this value is applied as a constant at the edge of the milling cutter. The temperatures applied to other heat sources of the screw system shown in Fig.\u00a07. The figure shows that the deformations of the upper and lower ends of the screw are larger, while those of the middle part are smaller. The centers of the upper and lower ends of the screw are taken as the research object, and the mean value of the maximum and minimum deformations of the ends is approximated as the deformation of their centers. The simulation results from the heat coupling deformation of the upper and lower ends of the ball screw at different times are shown in Figs.\u00a08 and 9, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000934_tmag.2019.2900527-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000934_tmag.2019.2900527-Figure7-1.png",
        "caption": "Fig. 7. 3-D RNA model of the variable inductor. (a) Division of the core. (b) Conventional 3-D unit magnetic circuit. (c) Proposed 3-D unit magnetic circuit incorporating the play model.",
        "texts": [
            " Therefore, the level of magnetic saturation in the ring yoke can be regulated by the primary dc current idc; that is, the effective inductance of the secondary ac windings can be controlled by the primary dc current. Inductor Incorporating Play Model Fig. 6 shows the specifications of the three-phase-laminatedcore variable inductor used for consideration. The core material is non-oriented silicon steel with a thickness of 0.35 mm. The rated capacity and voltage are 4.0 kVA and 200 V, respectively. The working flux density in the legs is 1.2 T. In the following, a method for deriving the RNA model is described. First, the variable inductor is divided into multiple elements, as shown in Fig. 7(a). To allow consideration of leakage flux, the surrounding space is also divided. Each divided element can be expressed in a 3-D unit magnetic circuit. Fig. 7(b) and (c) shows the conventional and proposed 3-D unit magnetic circuits in the core region. In the conventional RNA model, Rmr and Rm\u03b8 , which are the reluctances in the rolling plane, are determined in consideration of only the non-linear magnetic characteristics [11]. Hence, the magnetic hysteresis behavior is not taken into consideration. Therefore, in this paper, it is given by the play model and the circuit elements representing the classical eddy current loss and the anomalous eddy current loss, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002770_j.jmapro.2021.02.015-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002770_j.jmapro.2021.02.015-Figure2-1.png",
        "caption": "Fig. 2. The model of the SLMed circle structure with 90\u25e6 rotate scanning strategy: (a) layered by 90\u25e6 rotate; (b) overhead view.",
        "texts": [
            " When the \u03b1n is 90\u25e6, the minimum Rne and maximum Rni can be obtained. Due to the wall thickness Dnw is the difference between the Rne and Rni, so that the Dnw decreases first and then increases. In particular, if the \u03b1n increases continuously, the shape of the external and internal contours of the circle structure are oval-shaped approximately. Based on the fabrication manner of the layer by layer in SLM and the single-layer model, the model of the SLMed circle structure with 90\u25e6 rotate scanning strategy can be established, as shown in Fig. 2. As can be seen in Fig. 2, the most serious step effect occurs at the position 1/2/3/4, due to the first or the last track occur at these positions by turns. As a result, the circle structure with 90\u25e6 rotate scanning strategy has a bad surface quality, and its shape accuracy is distorted seriously. According to the Eqs. (3) and (4), the change rules of the geometric features fabricated by 90\u25e6 rotate scanning strategy can be seen in Table 2. As can L. Zhang et al. Journal of Manufacturing Processes 64 (2021) 907\u2013915 be seen in Table 2, as the \u03b1n increases from 0\u25e6 to 45\u25e6 and 90\u25e6 to 135\u25e6, the Rne and Dnw decreases, while the Rni increases; as the \u03b1n increases from 45\u25e6 to 95\u25e6 and 135\u25e6 to 180\u25e6, the Rne and Dnw increases, while the Rni decreases"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure18-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure18-1.png",
        "caption": "Fig. 18. Physical prototype of 3-R(RRR)R+R HAM.",
        "texts": [
            " 16 (a), (b) and (c), when the azimuth range of 3-R(RRR)R+R HAM is 0\u2013360\u25e6 continuous rotation and the pitch angle is 0\u2013390\u25e6 continuous rotation, the 3-R(RRR)R+R HAM can achieve the satellite tracking under the antenna pitch and azimuth motion. A prototype of 3-R(RRR)R+R HAM is made with a ratio of 1:2 to verify the mobility capability of the HAM. The virtual prototype model and explosion diagram of 3-R(RRR)R+R HAM are shown in Fig. 17. From the virtual prototype model and explosion diagram, the structure and assembly relationship can be shown clearly. Based on the virtual prototype model of the 3-R(RRR)R+R HAM, the parts of the HAM are processed and assembled. The physical prototype of the 3-R(RRR)R+R HAM is shown in Fig. 18. The motion experiment of 3-R(RRR)R+R HAM prototype is carried out to verify the motion performance of 3-R(RRR)R+R HAM. To facilitate the observation and removal of the antenna reflector connected to the polarization rotation mechanism, for the 3-R(RRR)R+R HAM prototype, 90\u25e6 pitch motion and 45\u25e6 pitch angle 360\u25e6 azimuth motion are, respectively, carried out and the motion experimental results are shown in Fig. 19. https://doi.org/10.1017/S0263574720001290 Downloaded from https://www.cambridge.org/core"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001357_j.oceaneng.2019.106812-Figure22-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001357_j.oceaneng.2019.106812-Figure22-1.png",
        "caption": "Fig. 22. Contour plotting of highest temperature during the rack and cylinder welding and bead-on-plate welding.",
        "texts": [
            " A new quarter FE model of cylindrical leg structure with bead-onplate welding on inner surface as shown in Fig. 21 was created and examined. The welding conditions as summarized in Table 1, temperature dependent material properties of EH36 as shown in Fig. 13, welding procedure and symmetrical boundary condition are all identical with the TEP FE computation above mentioned. By means of thermal analysis and mechanical analysis with considering the bead-on-plate welding on inner surface of cylinder, Fig. 22 shows the contour plotting of highest temperature during the actual welding and bead-on-plate welding. It can be seen that bead-on-plate welding on inner surface created a shallow welded zone and temperature distribution near the inner surface, which also cannot change the perimeter of examined cylinder but can reduce the influence of welding deformation during rack and cylinder welding on the cross section shape of cylindrical leg structure. Furthermore, when a number of 30 bead-on-plate welding passes on inner surface and directly under the rack welding region was practiced, Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001676_lra.2020.2996585-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001676_lra.2020.2996585-Figure2-1.png",
        "caption": "Fig. 2. Machine design of quasi-passive dynamic walker.",
        "texts": [
            " This section introduces the passive dynamics of this walker, as a prior knowledge for the following sections. As illustrated in Fig. 1, the position of grounding point on the hind side is (x, z) [m], the angular positions of the support and swing legs to the vertical are \u03b81 [rad] and \u03b82 [rad], respectively. Besides, the displacement of the wobbling mass with respect to the center is lc [m]. In addition, the body frame is always in parallel with the ground, due to the identicality, symmetry and synchrony between the fore and hind legs. Fig. 2 shows the detailed mechanisms of the legs synchronization realized by the connecting rods, and the mass wobbling motion realized by mapping it to the rotation of a DC motor via piston crank linkage. The generalized coordinate of the system is: q = [x z \u03b81 \u03b82 lc ]T. The walker is placed on a gentle slope with the angular position of \u03c6 [rad], and its swing legs generate a pendulum-like motion autonomously according to the following equation: Mq\u0308 + h = JT\u03bb. (1) Here,M is the inertia matrix,h is the combination of centrifugal force, Coriolis force and gravity terms",
            " It is however more robust when a slight perturbation, which is inevitable in the real-world, is added to the system. Moreover, we are taking step motion, that is, stance-leg exchange into account. Our future work will consider enlargement of the basin of attraction of the limit cycle to further enhance the stability [26]. Kinetic energy consumed by the ground collisions will be minimized according to the spring mechanism mentioned above. Moreover, experimental studies of a real machine according to Fig. 2 shall be also conducted and reported soon. REFERENCES [1] L. Amy, \u201cMake robot motions natural,\u201d Nature, vol. 565. pp. 422\u2013424. 2019. [2] ATLAS, \u201cBoston dynamics: Humanoid robot ATLAS,\u201d [Online]. Available: https://www.bostondynamics.com/atlas/ [3] M. Hutter et al., \u201cANYmal-a highly mobile and dynamic quadrupedal robot,\u201d in Proc. IEEE/RSJ Int. Conf. Intell. Robots Syst., 2016, pp. 38\u201344. [4] G. Bledt, M. J. Powell, B. Katz, J. Di Carlo, P. M. Wensing, and S. Kim, \u201cMIT Cheetah 3: Design and control of a robust, dynamic quadruped robot,\u201d in Proc"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002318_1729881420921644-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002318_1729881420921644-Figure2-1.png",
        "caption": "Figure 2. The schematic diagram of compliance errors caused by gravity.",
        "texts": [
            " The first method is clamping all the joints except the one to be measured and the second one measures the displacements of the EE due to certain forces and evaluates K throughout its Cartesian workspace. The latter method is more convenient and can provide a good approximation when the experiments are enough, which is adopted in this article. The solution formula for K is as follows14 F \u00bc J T KqJ 1Dxt \u00f013\u00de where F is the load on the EE, Dxt is the displacements of the EE and Kq is the 6 6 diagonal stiffness matrix. The numerical calculation method of equation (13) can refer to the literature.14 As shown in Figure 2, G and L represent the barycenter and length of the robot link, respectively, and q2 and q3 are the link angles relative to the vertical and horizontal lines. Then according to equation (12), the torsion angle of j2 caused by the gravity can be calculated by Dqc2 \u00bc C2 G2L2=2\u00fe G3L2\u00f0 \u00de sin q2 \u00fe C2G3L3 cos q3=2 \u00f014\u00de By the same token, the torsion angle of j3 is obtained by Dqc3 \u00bc C3G3L3 cos q3=2 \u00f015\u00de Geometric errors decoupled from compliance errors It can be known from equation (8) that the overall joint parameter errors are the superposition of geometric errors and compliance errors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.46-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.46-1.png",
        "caption": "Fig. 10.46 The structure of scraper discharge centrifuge [2]",
        "texts": [
            " The NC pulp in the storage tank 1 is conveyed into the flow tank 4 by the pump 2. After flow through the cleaner 3 and iron remover 5, NC is transferred to concentrator 6. The concentrated and heated NC slurry is pumped to the elevated tank 7, followed by the dehydration in the centrifugal dehydrator 8. After the dehydration, NC is packaged in bags for storage [2]. 2. The structure of centrifuge and dehydration process conditions \u2460 Scraper discharge centrifuge The structure of scraper discharge centrifuge is shown in Fig. 10.46. This equipment is continuously operated with automatic discharge, in-batch operation, high production capacity, and automatic feeding, dehydration, and discharge. However, it suffers some drawbacks, including complicated equipment structure, difficulty in controlling the operation, periodical cleaning maintenance of screen frame for improving efficiency, high labor intensity, and poor working conditions. Scraper discharge centrifuge consists of centrifuge spindle, bearings, a discharge device, loading device, sieve, chassis, filling, moving wheel, brake, and other components"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000460_s1560354716010081-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000460_s1560354716010081-Figure1-1.png",
        "caption": "Fig. 1. The Chaplygin sleigh on a cylinder.",
        "texts": [
            " In the case of a balanced and dynamically symmetric sleigh moving in a gravitational field it is shown that on an average the system has no drift along the vertical, and the motion of the sleigh is bounded by two horizontal planes (a similar situation takes place in a related problem due to Stu\u0308bler [14]). The system treated in this paper and similar systems arise in applied problems (for example, on the inner and outer surfaces of pipelines [15]) involving the motion of mobile robots on curved surfaces. 2. EQUATIONS OF MOTION The Chaplygin sleigh [5] is a rigid body with a sharp weightless wheel in contact with the (supporting) surface (see Fig. 1a). The sharp edge of the wheel prevents the wheel from sliding in the direction perpendicular to its plane. Let us consider the motion of the Chaplygin sleigh on a circular cylinder. We choose two coordinate systems: \u2014 an inertial coordinate system OXY Z in which the axis OZ coincides with the cylinder\u2019s axis (see Fig. 1b); \u2014 a noninertial (moving) coordinate system Px1x2x3 attached to the Chaplygin sleigh, with origin P at the point of contact of the wheel with the cylinder. Let us direct the axis Px1 along the normal to the cylinder\u2019s surface, and the axis Px2 along the plane of the wheel. We also assume that the sleigh moves so that the wheel\u2019s axis remains parallel to the tangent plane to the cylinder at the point of contact, P , and that the vector c from the point of contact P to the center of mass C remains constant in the system Px1x2x3",
            " Let R = (X,Y,Z) be the radius vector of the point of contact P in the coordinate system OXY Z and let (\u03c8, z) be its (dimensionless) cylindrical coordinates X = a cos \u03c8, Y = a sin \u03c8, Z = az, 3)We also note that the two above-mentioned papers by S.A.Chaplygin were republished in the English language in the journal Regular and Chaotic Dynamics. REGULAR AND CHAOTIC DYNAMICS Vol. 21 No. 1 2016 where a is the radius of the cylinder and \u03c8 \u2208 [0, 2\u03c0), z \u2208 (\u2212\u221e,\u221e). We define the rotation of the sleigh relative to the normal Px1 by the angle \u03d5 \u2208 [0, 2\u03c0) formed by the axis Px2 and the tangent to the level line z = const (see Fig. 1b). Thus, the configuration space N is a direct product: N = {q = (\u03d5,\u03c8, z), \u03d5, \u03c8 mod 2\u03c0} \u2248 T 2 \u00d7 R 1. The constraint equation in these variables becomes \u2212\u03c8\u0307 sin \u03d5 + z\u0307 cos \u03d5 = 0 (this constraint coincides with the Chaplygin constraint on the development of the cylinder on a plane [5]). The Lagrangian of the system has the standard form L = T (q, q\u0307) \u2212 U(q), where U(q) is the potential energy of external forces and T is the kinetic energy of the sleigh, which is represented as T = 1 2 mv2 \u2212 m(v, c \u00d7 \u03c9) + 1 2 (\u03c9, I\u03c9), (2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001631_j.promfg.2020.04.156-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001631_j.promfg.2020.04.156-Figure1-1.png",
        "caption": "Fig. 1. (a) Close-up of the laser head and (b) schematic of the coaxial laser beam and feeding lance arrangement.",
        "texts": [
            " The purpose of the present study is to establish an LMD wire process and to investigate the influence of variations of the interlayer dwell time on the related changes in microstructure. The paper is structured as follows: Section 2 gives an overview of the manufacturing of the samples from Ti-6Al-4V using the coaxial direct diode laser head with different dwell times. Also, the procedures of microstructural analysis are presented. Section 3 details the results and discussion of metallographic examinations. Finally, conclusions and outlook are presented. The LMD process was performed with a 1.2 kW direct diode laser head (Fig. 1) which was mounted on a Fanuc 6- axis robot (Fig. 2). Argon shielding gas was used to prevent the oxidation of the material. A beam transport fiber is not needed in this set-up. Only current and cooling water need to be connected to the LMD head. The filler metal in the form of wire can easily be fed perpendicular to the substrate surface into the laser spot. The most widely used titanium alloy Ti\u20136Al\u20134V was selected for the investigations in the present study. The geometry used in the current study represents a wall with a length of 98 mm and a width of 12 mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001617_b978-0-12-817450-0.00007-9-Figure7.3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001617_b978-0-12-817450-0.00007-9-Figure7.3-1.png",
        "caption": "Figure 7.3 Finite state machine of the neuromuscular simulation model to apply the FMCA-based control scheme that modulates the human ankle torque during walking.",
        "texts": [
            " Instead of using time-based patterns in the common trajectory tracking methods, here the GRF is employed as the phasing variable. Therefore, giving the FMCA-generated torque patterns to the prosthetic foot, releases dependency to time in trajectory-based methods. This force feedback includes other information (e.g., in case of perturbation occurrence), which is absent in methods using time-based trajectories. The simulation presented here uses the neuromuscular model including the prosthetic foot for one leg, developed by Thatte et al. [118]. Fig. 7.3 illustrates the finite state machine, defining the guard conditions for the transition between states within the complete gait cycle, introduced in Ref. [118]. This figure depicts the switching conditions between CP, CD, PP, and swing phase, while the control parameters are optimized for each phase. In the original simulation model, level transfemoral amputee walking was controlled by the reflex-based method [118]. We implemented the identified FMCA model from human walking on the neuromuscular simulation model, to control the ankle joint of a prosthetic foot"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003900_jsvi.1997.1323-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003900_jsvi.1997.1323-Figure3-1.png",
        "caption": "Figure 3. Configuration of the active bearing.",
        "texts": [
            " Finally, the matrices were condensed to a final dimension containing only a few normal co-ordinates. After the condensed mass and stiffness matrices are obtained, the equations of motion of the flexible sleeve can be adopted as follows: Ms \u00b7 r\u0308s +Ks \u00b7 rs =Hs +Cs , (1) where rs is the retained co-ordinates, Ms and Ks stand for the condensed mass and stiffness matrices, and Hs and Cs represent the hydrodynamic forces due to the instantaneous oil film pressure p and the chamber pressure pc respectively. The configuration of the active bearing is shown in Figure 3. The pressure distribution of the oil film is a function of the instantaneous displacements and velocities of the flexible sleeve and the journal, as well as the rotating speed of the rotor, i.e., p= p(V, rs , q, r\u0307s , q\u0307). The Reynolds equation was used to model the pressure distribution: 1 R2 1 18 0h3 h 1p 181+ 1 1z 0h3 h 1p 1z1=6V 1h 18 +12 1h 1t . (2) The thickness of the oil film can be written as: h= h'+Dh, (3) where Dh is the deformation of the sleeve which can be obtained by solving equation (1)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001811_s12008-020-00670-z-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001811_s12008-020-00670-z-Figure7-1.png",
        "caption": "Fig. 7 Reference frames positioned in each joint of a 4-DoF manipulator; a the reference frames are defined at the incorrect position; b the reference frames are defined at the correct position",
        "texts": [
            " 5 Blocks inside the Transformation A1, A2, and A4 blocks Fig. 6 Blocks inside the Transformation A3 block 1 3 transformation blocks A1, A2, and A4 (see Fig.\u00a04); the joint 3 performs a translational movement, so the \u201ctranslational variables axis z block\u201d must be selected inside the transformation block A3 (see Fig.\u00a06). The platform allows the user to improve their skills to position the reference systems in each joint of the robot (an example of a 4-DoF robot with the reference systems is shown in Fig.\u00a07b). When students load a robot model, all the reference systems are defined at the base of the robot (see Fig.\u00a07a); a student must change the definition of each reference system and visualize the movements that the robot performs. Following the rules of the DH method, students must define each reference system in the correct position (see Fig.\u00a07b); if it is necessary, the student can add more reference systems. For example, the DH method establishes that each z-axis is aligned with respect to the axis of movement of the joint by applying the corresponding transformations (always with respect to the previous system). For example, the student can analyze that in order to align the z-axis of system#1 with respect to the rotation system #2 of the robot (see Fig.\u00a07a), the z-axis must be rotated 90\u00b0 with respect to the x-axis of the system #0. Therefore, in the transformation A1 block (see Fig.\u00a05), it will modify in the rotation x-axis block by placing a value of 90, obtaining the new orientation shown in Fig.\u00a07b. Using problem-solving lessons is a typical teaching method used in undergraduate courses. The aim of the problemsolving lessons is to develop students\u2019 skills to select and deploy their theoretical knowledge to solve problems. Problem-solving lessons can be used to introduce new concepts and/or improve previously learned knowledge. In this way, students make sense of theoretical concepts by applying them to a variety of scenarios. Jonassen [41] discusses that the use of technology in teaching can be considered a pedagogical and cognitive tool that enhances the learning process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001627_icrom48714.2019.9071917-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001627_icrom48714.2019.9071917-Figure9-1.png",
        "caption": "Fig. 9: Wind tunnel test setup.",
        "texts": [
            " In this project, the mass parameters of the drone are obtained by 3D modeling, and the aerodynamic parameters are found by wind tunnel tests. For the motor blades system identification, a thrust, rpm, Pulse Width Modulation (PWM) measurement setup is used, and the relations between the PWM-Voltage-Thrust and the PWM-Voltage-rpm is identified. The tests results are shown in Fig. 8. These results are used as a look up table in the simulator to find the thrusts and rpm of the quadcopter\u2019s motor blade systems. For the VRS region identification, a wind tunnel tests was performed that is shown in Fig. 9. The last step for the MBD in the simulation phase is to find a proper controller for desired system. Based on the simulator, and the identified parameters, we found a controller for the Mambo drone to track a helix trajectory. The simulation results are illustrated in Fig. 10. Finally we used the matlab code generation to upload the controller codes on the Mambo drone. In the next section we will perform some flight test to find the behavior of the drone in the real flight situation. After uploading the code on the Parrot mini-quadcopter Mambo, it is better to do some flight in the External mode which can adjust the controller gains during the flight"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002221_0278364919897134-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002221_0278364919897134-Figure14-1.png",
        "caption": "Fig. 14. Test setup for external loading test. A 5.84 kg block was lifted up to a brace by the test sample. The inflation pressure was recorded while a photo was taken to show the differences in loaded geometry.",
        "texts": [
            " In addition to the coefficient of determination, the mean absolute error (in degrees) is also calculated MAE= 1 n Xn i = 1 ui Yij j \u00f029\u00de While the inflation test proves our ability to alter the deformation space of the elastomer membrane, it does not reveal how the membrane will behave when placed under a load. This load could come from an external force or from the fluid used to inflate the membrane. Here, we seek to determine the behavior of both the unreinforced and spherical fiber-reinforced membranes under both of these loads. 4.3.1. External loading. An external load was exerted on the inflated membrane perpendicular to the clamp base by a 5.84 kg aluminum block, shown in Figure 14. The membrane was inflated to the point where it pushed the block against a stabilizing fixture. The internal pressure of the membrane at this point was recorded and a picture was taken for analysis. Both membranes were also inflated to the same hemispherical shape before being loaded by the same aluminum block to directly compare their shape change. 4.3.2. Internal loading. An internal load was exerted on the membrane tangent to the clamp base by inflating the membrane with water. Images were taken at various pressures to compare the loaded membranes\u2019 geometry"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure21.5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure21.5-1.png",
        "caption": "Fig. 21.5 Relative arrangement of the LGC for the operator in the lower extremities exoskeleton (1 the foot, 2 the lower limb, 3 the thigh, 4 the back, 5 the linear gravity compensator)",
        "texts": [
            " The combined LGC operation principle is that, the potential energy of the elastic element accumulates during its\u2019 compression, and then, when it is stretched, the energy is transmitted to the exoskeleton in the form of a moment MO43(\u03a61) applied by the HJ. In parallel with this, the electric drive installed in LGC creates a moment MO43(\u03a62). Thus, we get an assisting moment acting on the exoskeleton\u2019s back 4 relative to the hip 3 and helping the person to perform the necessary movements. MO43(\u03a6) = MO43(\u03a61) + MO43(\u03a62). (21.1) Let us consider in more detail the HMI structural diagram, which provides the interaction of a human andBTWS exoskeleton, shown in Fig. 21.5. TheHMI consists of a back module that includes the exoskeleton\u2019s and the operator\u2019s back absolute inclination angle sensors, a force interaction sensor, a hip joint equipped with a relative angle meter, an on-board computer, and LGC HJ. By analyzing the information received from the sensors in accordance with the developed algorithm, the on-board computer generates control voltages supplied to the LGC electric drive. For the development and research of the BTWS, a kinematic and dynamic model of the system has been developed that allows us to work out in detail the process of interaction of a biological object, the electromechanical part of the exoskeletal system, and HMI",
            " The operator sets the movement nature of the exoskeleton\u2019s back using the back inclination angle \u03c65. The considered systemhas two generalized coordinates\u03c64 and\u03c65. The position of the operator\u2019s back sets the angle \u03c65, and the position of the exoskeleton\u2019s back sets the angle \u03c64. It is necessary to connect the coordinates of points A, B, which determine the exoskeleton\u2019s kinematics and the generalized coordinate \u03c64 between themselves. For this purpose, it is necessary to consider the circuit shown in Fig. 21.5. To solve the problem of determining the relationship between the characterical system points corresponding coordinates, the vector-matrix method can be applied. We write the equality: r B = r A + r AB . (21.2) From here, the distance between points A and B corresponding to the length L of the LGC can be found. r AB = r B \u2212 r A (21.3) or, taking into account (21.2) (21.3), can be got: r AB = r4A \u2212 r3B, (21.4) where 21 Modeling of the Exoskeletal Human-Machine \u2026 265 r AB = \u2223 \u2223 \u2223 \u2223 xAB yAB \u2223 \u2223 \u2223 \u2223 ; r3B = \u2223 \u2223 \u2223 \u2223 xO43B yO43B \u2223 \u2223 \u2223 \u2223 ; r4A = \u2223 \u2223 \u2223 \u2223 xO43A yO43A \u2223 \u2223 \u2223 \u2223 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001156_j.conengprac.2019.104118-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001156_j.conengprac.2019.104118-Figure2-1.png",
        "caption": "Fig. 2. Visualization of the investigated system. The whole system and the energy recovery system are depicted. The ERC-system is marked red, whereas the hydraulically actuated cylinders are marked green. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " The investigated system is a material handling excavator. The excavator consists of a lower structure, an upper structure, a boom and a stick. Usually, a tool is attached to the stick which is neglected for the remainder of this paper. The equipment (boom and stick) is actuated by hydraulic cylinders. The boom has an additional gas filled cylinder for energy recovery, which is called ERC-system. The energy recovery cylinder is equipped with a cooling and heating circuit to keep the gas at an operating temperature. Fig. 2 shows the whole system and the ERC-system in particular. The hydraulic cylinders are marked green and the energy recovery cylinder is marked red within Fig. 2a. Note the different installation positions of the hydraulic cylinders and the gas-filled cylinder on the boom, see. Fig. 2b, which results in different kinematic relationships of the cylinder length and position of the boom. The hydraulic system consists of two variable displacement pumps for the equipment, valves for each equipment part and is a loadsensing system type. The load-sensing system uses a load independent flow distribution (LUDV) technology which results in load independent stationary stroke velocities of the hydraulic cylinders even in the case of a saturation condition. The LUDV is only active for lifting motions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003855_jmbi.1995.0572-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003855_jmbi.1995.0572-Figure6-1.png",
        "caption": "Figure 6. Assignment of the subunit domains to the sequence. (a) Contoured map of the flagellar filament in axial view, with the four domains of each subunit labelled from D0 to D3. The map was calculated with the data obtained by Mimori et al. (1995), but by limiting the resolution to 12 A\u030a. The thickness of the map is 50 A\u030a, which contains 11 subunits in two turns of the one-start helix. Two neighboring subunits along the left-handed five-start helix are indicated. The subunit boundary lines in the core region (broken lines) are drawn using radial lines, because the boundary in this region cannot be identified unambiguously at this resolution. The scale bar represents 100 A\u030a. (b) Positions on the sequence of the proteolytic fragments, predicted secondary structures and assigned domains. The positions of the two major proteolytic fragments are indicated with F27 and F40. Predicted secondary structures are: open box, a-helix; closed box, b-sheet; small open box, turn (Vonderviszt et al., 1990). The shaded bars represent the conserved a-helical coiled-coil forming regions (Figure 7(a)).",
        "texts": [
            " In contrast to the core region, the densities from about 60 to 80 A\u030a are relatively low and those from 80 to 120 A\u030a been identified in the fragments obtained by limited proteolysis. A similar interpretation has been made by Mimori et al. (1995) based on the volume fractions of the domains and the Discussion below is consistent with it and confirms it. The structure factors obtained by Mimori et al. (1995) were used to calculate a density map of the 50 A\u030a thick section by limiting the resolution to 12 A\u030a, so that the map does not show too much detail in the subunit shape. The map is shown in Figure 6(a), in which two neighboring subunits along the left-handed five-start helix are indicated and the domains are labelled as in Mimori et al. (1995): the inner tube, D0; the outer tube, D1; the vertical column, D2; the outer domain, D3. Subunit boundary lines in the core region (the domains D0 and D1) are artificially set with radial lines, because it is not possible to unambiguously determine the boundary in this high density region. The whole sequence of 494 amino acids is divided into two parts, the terminal regions, from Ala1 to Arg65 and from Ser451 to Arg494, which have no ordered tertiary structure in the monomeric state (Vonderviszt et al",
            "9 kDa, named F27, which consists of two domains as deduced from the deconvolution analysis of the calorimetric melting profile (Vonderviszt et al., 1990), and the remainder forms the third domain of F40, composed of two discontinuous sequences from Asn66 to Lys177 and from Ser423 to Arg450 (Vonderviszt et al., 1989). The numbering given here is based on the correct sequence published afterwards by Kanto et al. (1991), and therefore slightly different from the original one of Vonderviszt et al. (1989). In Figure 6(b), the sequence positions of these fragments are indicated on the line representing the flagellin sequence, on which the positions of predicted secondary structures are also indicated. F27 contains the sequence known to be exposed on the filament surface (Trachtenberg & DeRosier, 1988; Fedorov et al., 1988). The mass of the outer part, namely D2 and D3 together, corresponds to about 24 kDa as calculated from the mass fraction in Table 1, and this is close to the molecular mass of F27, 24.9 kDa. This suggests that the domains D2 and D3 are the two domains of F27. The third domain of F40, whose molecular mass is about 15.3 kDa, can be tentatively assigned to a major part of the domain D1, the outer tube, of about 20 kDa, because the third domain is closely associated with F27, together forming the compact F40, and the domain Dl has the closest association with the domain D2 while being rather separated from the domain D0 with the gap shown in Figure 6(a). Then the mass of the domain D0, about 8 kDa, is assigned to cover about two thirds of the terminal regions of 11.5 kDa. Both terminal regions seem to make almost equal contributions to the domain D0, because the reconstituted filaments from terminally sets of phasing procedure (Figure 3), in which all the data sets produced the consistent phase assignment. The radial mass distribution so deduced showed only a small difference from the radial volume distribution estimated from the density map of helical image reconstruction (Table 1)",
            " The mass distribution will be correlated with the domains (see the Discussion below), to make a more quantitative assignment of the morphological domains to specific parts of the sequence that have and of those seven a-helices the two closest to both termini are in the domain D0 and the rest five in D1. Coiled-coil prediction based on evolutionary information was applied to reveal which a-helical segments contribute to the conserved coiled-coil features. As clearly seen in Figure 7(a) (also indicated in Figure 6(b)), the conserved helicalbundle forming regions are preferentially located in the third domain of F40. This result supports our domain assignment described in Figure 6, in which these regions are assigned within the domain D1, the outer tube of high densities. It is also consistent with the density map of the outer tube, containing many rod-like features with near-axial orientation, which are likely to represent a-helical bundles (Mimori et al., 1995; Morgan et al., 1995). Coiled-coil prediction specific for Salmonella typhimurium sequence is shown in Figure 7(b) and it suggests that the coiled-coil regions are further extended in the flagellar filament we studied",
            " These two binding positions are in the outer part (the domains D2 and D3), where the subunit boundaries are clearly identified and the neighboring subunits make contacts along the left-handed five-start helix (Mimori et al., 1995). Those divalent or trivalent cations are likely to be bound at the interface between the subunits, enforcing the interaction and stabilizing the filament structure. The mercury binding position in the filament of mutated flagellin, G365C, indicates that the radial position of Gly365 is close to 61 A\u030a from the filament axis, which is in the domain D2. Gly365 is in the COOH-terminal half of F27 as shown in Figure 6(b). Recently, other cysteine residues were introduced in this region and they were also located in D2 (data not shown). This suggests that the COOH-terminal half of F27 forms the domain D2 and the NH2-terminal half forms D3. truncated fragments show no clear staining in the central part of the filament in electron micrographs of negatively stained specimens, unless at least 30 residues are removed from both terminals of flagellin (Vonderviszt et al., 1991). A preliminary map by electron cryomicroscopy of these filaments in the frozen hydrated state has also shown the almost complete disappearance of the inner tube by deleting about 30 residues from both terminals while the outer tube remains intact (Y.M., F.V. & K.N., unpublished results). Therefore, the two predicted a-helices closest to the termini are in the domain D0 and most of the other a-helices are in D1. These assignments are summarized in Figure 6(b). Position of a-helical coiled-coils As described by Namba et al. (1989), X-ray fiber diffraction patterns of the flagellar filaments show strong layer-lines around 1/5 A\u030a\u22121, which indicates the existence of a-helices that are aligned parallel to the filament axis. Near-equatorial intensity distribution is strong around 1/10 A\u030a\u22121, and cylindrical Patterson analysis of the intensity distribution indicates that these a-helices form coiled-coils in the core region of the filament. The a-helical bundles are supposed to play an essential role in flagellar filament formation (Homma et al., 1990). There are four a-helices predicted in the terminal regions that are disordered in the monomeric state and three in the third domain of F40 (Vonderviszt et al., 1990). As shown in Figure 6(b), the core region of the filament, the domain D0 and D1 in Figure 6(a), are assigned to cover both terminal regions Folding pattern of flagellin subunit and its implications Based on the domain-sequence assignment and other information described above, the folding pattern of flagellin can be roughly described as follows. The NH2-terminal starts from the domain D0, after about 30 to 40 residues the main chain goes out to the domain D1 and folds mainly into a-helical coiled-coil. Then from around Tyr178 the chain goes out to the outer part, probably the domain D3 first and then into D2, folding into mainly b-sheet until it reaches to around Arg422"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001241_j.conengprac.2019.104207-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001241_j.conengprac.2019.104207-Figure4-1.png",
        "caption": "Fig. 4. The lab-scale overhead crane experimental set-up. The relevant components are highlighted in the figure. A view from the top details the considered direction of movement of the crane.",
        "texts": [
            " Note that, from practical perspective, relative degree can change with respect to time due to change of working point or change in the system configuration. Relevant examples can be a micro-grid system, which exhibits relative degree equal to one in the grid-connected operation mode, whilst relative degree is equal to two in islanded operation mode (Cucuzzella, Incremona, & Ferrara, 2015). Also the kinematic dynamics of vehicle motion exhibits changes in relative degree (Levant & Livne, 2016). The lab-scale overhead crane shown in Fig. 4 is employed to experimentally assess the relative degree identification scheme presented in Section 3. Even though the crane can move in all the three dimensions (\ud835\udc65 \u2212 \ud835\udc66 \u2212 \ud835\udc67 axis in Fig. 4), only the \ud835\udc65-displacement is considered in this framework. This because the other two displacements display similar features. The experimental set-up is available at Process Control Laboratory of University of Pavia, and it is composed by the following key-elements: (\ud835\udc56) the structure of the lab-scale overhead crane; (\ud835\udc56\ud835\udc56) a 12 Volts DC motor, mechanically coupled with the transmission belt in Fig. 4; (\ud835\udc56\ud835\udc56\ud835\udc56) the encoder (CUI AMT-102V), which is fixed to the belt shaft as shown in Fig. 4. The encoder is capable of measuring the angular position of the shaft of the belt, which is related to the \ud835\udc65- displacement of the crane by making use of basic mechanical relations. The accuracy of the encoder has been set in such a way that it is possible to track the movement of the crane with a resolution of 1 millimetre. The interested reader is referred to CUI (2018) for the technical data of the adopted encoder. The sampling time \ud835\udf0f for the measurement and for the test-input signal is chosen equal to \ud835\udf0f = 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003126_s42835-021-00852-z-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003126_s42835-021-00852-z-Figure1-1.png",
        "caption": "Fig. 1 The schematic\u00a0view of a 2\u00a0MW DD HTS generator",
        "texts": [
            " Direct-drive (DD) wind power generators with large power ratings will reduce the cost of transportation, installation, and foundation of wind turbines and consequently lower the Levelized Cost of energy [9], especially for offshore wind farms. Compared with traditional DD wind power generators, DD HTS generators could be a promising alternative because of their superior torque density and high efficiency [10]. Considering the design difficulty of the rotating current collector and the installation of cryogenic refrigeration system, a 2\u00a0MW DD HTS generator employ HTS wires in the rotor field windings and copper conductors in the stator coils were designed in this study, the schematic\u00a0view of the generator is shown in Fig.\u00a01. Especially, the rotor field windings wound by HTS wires can generate a much higher magnetic field than that of the conventional copper coils of the same size. To overcome the iron teeth saturation caused by high air-gap magnetic density, the iron teeth between stator coils can be replaced by composite teeth, such as glass-fiber * Yong Zhou wade_812@163.com 1 College of\u00a0Civil Engineering and\u00a0Mechanics, Lanzhou University, Lanzhou, China 2 Wuhan Institute of\u00a0Marine Electric Propulsion, China State Shipbuilding Corporation Limited, Wuhan, China 1 3 reinforced plastic (GFRP) teeth, i",
            " A 2\u00a0MW DD HTS wind power generator was designed and the electromagnetic (EM), EM force, loss, and insulation of the generator were analyzed and developed, respectively. The thermal and mechanical of the one pole pair stator was analyzed by the finite element analysis (FEA). A pole pair full-scale oil-cooling air-core stator was designed, manufactured, and tested. A 2\u00a0MW DD HTS generator is mainly composed of the oil-cooling air-core stator, HTS magnet, torque transfer system, coolant transfer coupling, stator cooling system, and cryogenic refrigeration system, as shown in Fig.\u00a01. The oilcooling air-core stator is consisted of back iron, GFRP teeth, windings, and a composite cylinder, as shown in Fig.\u00a02. Especially, the composite cylinder locates in the inner radial surface of the wedge for a more rigid assembly of the stator and is also used to isolate the stator and rotor and form a closed space with the stator frame for the flowing of the cooling oil, which is not shown in Fig.\u00a02. The flux density has a direct impact on the performance and cost of the HTS generator, and that of air-gap, rotor yoke, and stator yoke after comparative analysis are chosen to be 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003126_s42835-021-00852-z-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003126_s42835-021-00852-z-Figure8-1.png",
        "caption": "Fig. 8 Von-Mises stress contour of the slot region of the stator",
        "texts": [
            " The outer surface of the back iron is fixed and the two parallel symmetry faces could translate in the X and Y directions, while fix in the Z direction. The defined direction is similar as shown in Fig.\u00a03. The twelve conductors with two columns in the slot region are numbered 1 to 6 correspond with the EM modal. By loading the EM forces from the EM analysis (results summarized in Table\u00a02) and the temperature distribution from the CFD analysis (as shown in Fig.\u00a06). The Von-Mises stresses contour in the slot region of the stator is shown in Fig.\u00a08. The contour indicates that there is a stress concentration. The maximum stress is about 290\u00a0MPa and it is located in the bottom corners of the GFRP teeth because of the different thermal expansion coefficients between the copper conductor and the GFRP teeth. Especially, the tension strength of the GFRP is about 452\u00a0MPa [25]. Therefore, the safety factor is about 1.6 (452/290 = 1.6), then the GFRP teeth meet the strength requirement of the HTS generator. The resultant stresses in the wedge components are extremely low at approximately 15\u00a0MPa"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000255_icems.2019.8921846-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000255_icems.2019.8921846-Figure1-1.png",
        "caption": "Fig. 1. Radial and axial cross sections of the electrical machine.",
        "texts": [
            " LPTN MODELS AND CRITICAL ISSUES Since the LPTN model [4] proposed by Mellor in 1991, a variety of LPTN models have been developed, the accuracy of sophisticated models [4], [5] or simplified models [6]-[12] is closely dependent of the accuracy of some critical parameters, including: \u2022 Equivalent thermal resistance between the frame and the ambient; \u2022 Convection heat transfer coefficients within the end-cap; \u2022 Winding modelling and equivalent winding thermal conductivity. In order to investigate the above-mentioned aspects, the thermal analysis software Motor-CAD is employed in this D 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE paper. Fig. 1 presents the structure of the prototype 12- slot/10-pole IPMSM. Detailed information of this thermal network can be found in [22], [23]. The geometric parameters of the analyzed machine are presented in Table A.I in Appendix. Besides, the analytical LPTN models developed by various authors [4]-[15] provide an alternative approach to estimate temperatures of electrical machine. Compared to the software package, the application of analytical LPTN model is more flexible, i.e. the analytical thermal model can be easily embedded in further thermal analyses, and get rid of the limitations of electrical machine types, which have been prestored in the software"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000292_s11432-019-1470-x-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000292_s11432-019-1470-x-Figure2-1.png",
        "caption": "Figure 2 (Color online) Modules of the UGSR. (a) Structures of the rotate and telescopic modules; (b) composition of the telescopic module.",
        "texts": [
            " Section 4 derives the input-output linearization of the model, and obtains its normal form. The system control law designed by the SMC method, and the nonlinear observer designed by the UKF method, are also introduced in this section. Finally, the proposed controller and observer are verified by simulations in Section 5. This section describes the development and implementation of the UGSR prototype. Figure 1 shows the gliding motion and snake-like swimming of the robot during pool tests. The UGSR has a serial structure composed of multiple telescopic and rotate modules (see Figure 2). Each telescopic module has three DOFs, namely, telescoping, pitch, and yaw. The net buoyancy can be adjusted by controlling the length of the telescopic module. A pitch moment is generated when the telescopic modules symmetrically installed at both ends of the robot are unequal in length. The net buoyancy and pitch moment are coupled, as both are related to the elongations of the two telescopic modules. The UGSR is upwardly gliding in Figure 1(a). The rotate module has two DOFs, namely, pitch and yaw"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001152_ab3f56-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001152_ab3f56-Figure6-1.png",
        "caption": "Figure 6. Setup of test bench used to evaluate performance.",
        "texts": [
            " A moving average filter with a window size of 100 was applied to the measured value to reduce the noise effect. The control board uses an MCU (Teensy 3.6, PJRC, OR, USA) to control the tension of the pulling wire. This is achieved by changing the rotational velocity of the motor via a proportional\u2013integral\u2013derivative (PID) feedback control at 1 kHz based on the pressure sensor value. To evaluate the compression performance of the proposed sleeve, a test bench was constructed by placing a cylindrical inflatable tube on a rigid cylinder made of ABS, as shown in figure 6. To reproduce deformation of the body from the compression force of the sleeve, a cylindrical inflatable tube was made using polyvinyl chloride (PVC) fabric, and this was then deformed using the external compression force. The compression force of the sleeve was measured by the inflatable tube on the test bench. A pressure sensor was installed to the inlet hose of the inflatable tube to measure the changes in the internal air pressure. The pressure sensor was connected to the measurement board to read the data in real time"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003066_j.compeleceng.2021.107267-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003066_j.compeleceng.2021.107267-Figure14-1.png",
        "caption": "Fig. 14. Analysis through Finite Element Method (FEM), seeking to validate its performance against dropping conditions.",
        "texts": [
            " Hence, in order to summarize the most relevant results of the student in terms of the methodology, Fig. 13 shows an illustrative summary of the results documented from the final presentation of the student. Where, in the first point an example of the figures used for the concepts introduction is added, meanwhile the second and third points from Fig. 13 are from the design and system simulation performed by the student. Both stages (stage 2 and stage 3) can be complemented by the evidence shown in Fig. 14, where the implementation of simulation tools on the design proposed by the student can be observed in greater detail, where, prior to the implementation stage, the performance simulation was carried out of the prototype against possible falls and blows in the critical points of the drone. Moreover, the base of the V-Model from Fig. 13 shows the experimental quadcopter prototype the student deployed; which was later was tested and the controller outputs of the implemented fuzzy controller is shown in stage 5, in addition to the validation of the system that was performed by the student by plotting the error through time of the stabilization signal",
            " 5 and 13, the V-Model for education proposed in this work not only gives a framework in which a student\u2019s project can be designed, implemented and evaluated, but also give a clear structure of how the role of the teacher changes throughout the iterative process (as validated through Figs. 6 and 7), starting with greater importance at the beginning of the project where the concepts are presented, going through almost no participation in the advisory role where students take the leading role, until the ending point where the teacher takes the primary role again to validate and give feedback to the students on the development of the project. From Fig. 14, it is clear that the student could develop the designing skills of a mechanical structure in spite of the electronics engineering background of the testing subject, which clearly validates that the student learned new competences through the experimental development of the quadcopter prototype. Yet, there were some limitations in the mechanical design shown in Fig. 15, since the student\u2019s inexperience in the field, made him had some limitations on some finer design details, such as: the selection of the structural materials, the total weight of the prototype and finally its aerodynamics"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001759_j.mechmachtheory.2020.103995-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001759_j.mechmachtheory.2020.103995-Figure2-1.png",
        "caption": "Fig. 2. The double spherical 6 R linkage.",
        "texts": [
            " The intersections of creases z 1 and z 6 , z 5 and z \u2019 6 are denoted by S 1 and S \u2019 1 , respectively. Because of plane-symmetry, creases z 2 , z 3 and z 4 intersect at point S 2 with sector angles while \u03b161 = \u03b156 = \u03c0 \u2212 \u03b1. A close-loop kirigami linkage in Fig. 1 (b) is formed by joining panels p 1 and p 6 with z 6 and z\u2019 6 collinear as one crease. The geometry conditions of this linkage are \u03b123 = \u03b134 = \u03b1, \u03b156 = \u03b161 = \u03c0 \u2212 \u03b1, a 12 = a 45 = a. (1) Based on this kirigami pattern, an overconstrained 6 R linkage is derived in Fig. 2 , where the panels and creases of crease pattern are replaced by links and revolute joints of linkage, respectively. Following the D-H notation [32] , the axis z i is along the revolute joint i , x i is the common normal from z i \u22121 to z i , and y i is determined by the right-hand rule. a i ( i + 1) is the link length between axes z i and z i + 1 . \u03b1i ( i + 1) is the twist angle from axes z i to z i + 1 positively about x i + 1 . R i is the offset from x i to x i + 1 positively about z i . In addition, variable \u03b8 i is the angle from x i to x i + 1 along the positive direction of z i . For the plane-symmetric 6 R linkage, all the coordinates on the joints are setup accordingly in Fig. 2 , and the distances from the revolute joints to the centres of spheres S 1 and S 2 are r 1 and r 2 respectively. Considering the symmetry of the kirigami pattern in Fig. 1 , in the linkage in Fig. 2 , axes of three adjacent revolute joints 5, 6 and 1 intersect at spherical center S 1 , while the other three axes of revolute joints intersect at another spherical center S 2 , which reveals that the linkage belongs to the double spherical 6 R linkage. Furthermore, axes of joints 3 and 6 form the symmetry plane \u041f of the derived 6 R linkage. Hence, its geometric conditions are a 12 = a, a 23 = 0 , a 34 = 0 , a 45 = a, a 56 = 0 , a 61 = 0 \u03b112 = 0 , \u03b123 = \u03b1, \u03b134 = 2 \u03c0 \u2212 \u03b1, \u03b145 = 0 , \u03b156 = \u03c0 + \u03b1, \u03b161 = \u03c0 \u2212 \u03b1 R 1 = r 1 , R 2 = \u2212r 2 , R 3 = 0 , R 4 = r 2 , R 5 = \u2212r 1 , R 6 = 0 , (2) in which \u03b1 \u2208 (0 , \u03c0) ",
            " Its closure equation is T 21 T 32 T 43 = T 61 T 56 T 45 , (3) where T (i +1) i = \u23a1 \u23a2 \u23a3 cos \u03b8i \u2212cos \u03b1i (i +1) sin \u03b8i sin \u03b1i (i +1) sin \u03b8i a i (i +1) cos \u03b8i sin \u03b8i cos \u03b1i (i +1) cos \u03b8i \u2212sin \u03b1i (i +1) cos \u03b8i a i (i +1) sin \u03b8i 0 sin \u03b1i (i +1) cos \u03b1i (i +1) R i 0 0 0 1 \u23a4 \u23a5 \u23a6 , (4) and T i (i +1) = T \u22121 (i +1) i = \u23a1 \u23a2 \u23a3 cos \u03b8i sin \u03b8i 0 \u2212a i (i +1) \u2212 cos \u03b1i (i +1) sin \u03b8i cos \u03b1i (i +1) cos \u03b8i sin \u03b1i (i +1) \u2212R i sin \u03b1i (i +1) sin \u03b1i (i +1) sin \u03b8i \u2212 sin \u03b1i (i +1) cos \u03b8i cos \u03b1i (i +1) \u2212R i cos \u03b1i (i +1) 0 0 0 1 \u23a4 \u23a5 \u23a6 . (5) It is assumed that the linkage maintains plane symmetry in the process of motion. Due to the coordinate set-ups of in Fig. 2 , we have \u03b81 = \u03b85 , \u03b82 = \u03b84 . (6) Substituting Eq. (2) to Eq. (3) and then simplifying element (3, 3) of entries (shown in Appendix A), the following equa- tion is obtained. ( cos \u03b83 \u2212 cos \u03b86 )( cos 2 \u03b1 \u2212 1) = 0 , (7) which can be further simplified as \u03b83 = \u03b86 , (8) or \u03b83 = \u2212\u03b86 . (9) With the element (3, 4), we can get the relationship between kinematic variables \u03b81 and \u03b86 . tan \u03b86 2 = a cos \u03b81 ( r \u2212 r ) sin \u03b1 \u2212 a cos \u03b1 sin \u03b8 . (10) 2 1 1 If \u03b83 = \u03b86 , then simplified elements (1, 1) and (3, 1) can be written as cos \u03b1 sin \u03b83 sin ( \u03b81 + \u03b82 ) = 0 , (11) and sin \u03b83 + sin \u03b81 sin \u03b82 sin \u03b83 \u2212 cos \u03b1 cos \u03b81 sin \u03b82 + cos \u03b1 cos \u03b81 cos \u03b83 sin \u03b82 \u2212 cos \u03b81 cos \u03b82 sin \u03b83 \u2212 cos \u03b1 cos \u03b82 sin \u03b81 + cos \u03b1 cos \u03b82 cos \u03b83 sin \u03b81 = 0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000684_cjme.2016.0617.074-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000684_cjme.2016.0617.074-Figure3-1.png",
        "caption": "Fig. 3. Model of gas film thickness disturbance analysis",
        "texts": [
            " when the shaft spins, the installation deviation of rotor will provide an angular excitation motion to stator. The spiral groove geometry is shown in Fig. 2(a), when the rotor revolves at a high speed, there is a large hydrodynamic pressure will be generated by the spiral grooves, which will help to separate the two seal faces and maintain a steady gas film thickness. But in most of the practical seals, the gas film thickness has a disturbance because of the rotor\u2019s axial runout and misalignment. Fig. 3(a) shows the relative position between stator and rotor, Fig. 3(b) shows the model of seal face kinematics, and in the seal dynamic tracking property analysis, the gas film, which have a certain stiffness and damping properties, is usually treated as a spring and damper system, and the stator can be thought of as being supported by it[3]. Stator tracks rotor\u2019s excitation motions under the function of spring, o-ring and gas film, and a good dynamic tracking property can help the seal to continuously operate with no face contact or no excessive leakage. 2.2 Mathematical models Assuming the ideal gas and isothermal conditions, and without considering the impact of seal faces deformation in high-pressure"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002357_j.triboint.2020.106481-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002357_j.triboint.2020.106481-Figure2-1.png",
        "caption": "Fig. 2. Preload-Displacement.",
        "texts": [
            " [32] for the approximate calculation of Hertzian elliptic contact parameters given in Eqs. Eq. (3)\u2013(6) and elliptic integrals. \ud835\udc4e\u2217 = ( 2\ud835\udf052(\ud835\udf05) \ud835\udf0b )1\u22153 (3) \ud835\udc4f\u2217 = ( 2(\ud835\udf05) \ud835\udf0b\ud835\udf05 )1\u22153 (4) \ud835\udeff\u2217 = 2(\ud835\udf05) \ud835\udf0b ( \ud835\udf0b 2\ud835\udf052(\ud835\udf05) )1\u22153 (5) \ud835\udc58 = 4 \u221a 2\ud835\udc38\u2032 3(\ud835\udeff\u2217)3\u22152\ud835\udc45 (6) \ud835\udc39\u210e = \ud835\udc58\ud835\udeff3\u22152 (7) When the angular contact ball bearings are subjected to an axial preload, an equal load distribution and contact angle is formed in each rolling element. The contact angle in the preloaded case is greater than the free contact angle. The deformation \u2018\u2018\ud835\udeff0, \ud835\udc670\u2019\u2019 caused by the preload can be obtained by using Eqs. (8)\u2013(9) from Fig. 2. \ud835\udc35\ud835\udc51\ud835\udc4f((cos \ud835\udefc0\u2215 cos \ud835\udefc\ud835\udc5d)\u2212 1) term in Eq. (8) is equal to \ud835\udeff0. \ud835\udc43\ud835\udc5f \ud835\udc5a sin \ud835\udefc\ud835\udc5d = \ud835\udc58 ( \ud835\udc35\ud835\udc51\ud835\udc4f ( cos \ud835\udefc0 cos \ud835\udefc\ud835\udc5d \u2212 1 ))3\u22152 (8) \ud835\udc670 = \ud835\udc35\ud835\udc51\ud835\udc4f ( sin (\ud835\udefc\ud835\udc5d \u2212 \ud835\udefc0) cos \ud835\udefc\ud835\udc5d ) (9) 2.2. EHL Theory The mechanics of a moving fluid are studied by hydrodynamic theory. The hydrodynamic flow regime where the hydrodynamic pressure is great enough to form an elastic deformation of the solids in contact is studied by EHL theory. However, it is shown by many researchers that in the presence of EHL contact, the pressure distribution on the relative moving surfaces differs slightly from the Hertzian pressure, especially exit region of contact [11,28,33]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001594_j.cma.2020.112996-Figure17-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001594_j.cma.2020.112996-Figure17-1.png",
        "caption": "Fig. 17. Helical rod clamped at one end and subjected to a vertical load at the free end.",
        "texts": [
            " 16, one observes that the two new formulations are excellent in predicting the stress resultants. The large oscillation of the torque by the \u03c6Rd0 formulation occurs mainly in the bulged patch, while for the bending moment it produces incorrect result even in the straight patch near the clamped end. It should be noted this formulation is invariant only for straight and circular members, and the error is due to the incorrect end moments computed, caused by the bulging part. To test the performance of each formulation to cope with general 3-dimensional rods, a helical rod [5] shown in Fig. 17 is considered, of which the initial rod axis is r(\u03b2) = [ \u03c1 cos \u03b2 \u03c1 sin \u03b2 a 2\u03c0 \u03b2 ]T (74) where \u03c1 is the helix radius, the angle \u03b2 \u2208 [0, 2m\u03c0 ], m is the number of loops (with m = 5 adopted in this study). The tangent d1 (\u03b2) can be calculated as d1 (\u03b2) = [ \u2212 sin \u03b2 cos \u03b2 a 2\u03c0\u03c1 ]T /\u221a 1 + ( a 2\u03c0\u03c1 )2 (75) and the transversal vectors are defined as d2 (\u03b2) = [cos \u03b2 sin \u03b2 0]T and d3 (\u03b2) = d1 (\u03b2) \u00d7 d2 (\u03b2) (76a,b) The rod is clamped at one end, r(0), and subjected to a vertical load P = [0 0 Pz]T at the free end, r(2m\u03c0 )"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002542_icem49940.2020.9270965-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002542_icem49940.2020.9270965-Figure9-1.png",
        "caption": "Fig. 9. Schematic of modular in-wheel motor [25].",
        "texts": [
            " In the formal architecture, two separate stators with separate driven inverters are embedded into a common housing and rotor, adopting single layer FSCW configuration; the latter one, two set of three-phase in one stator with double-layer FSCW configuration driven by two independent inverters. The outputs of simulation and experimental tests showed the excellent fault tolerance. To ensure adequate levels of functional safety, a FT concept for the design of in-wheel motors has been demonstrated [25], in Fig. 9. Considered appropriate inductance to limit fault currents, the competence of FT can be achieved by a series of eight independent sub-motors. Experiment has shown a good magnetic isolation between sub-motors, so the failure of any sub-motor should not significantly affect the operation of others. 2) Flux Controllable PM Machine Besides the conventional rotor-PM machines, doublesalient PM (DSPM) machine, memory machine and flux switching PM (FSPM, in Fig. 10) machines have been investigated[29, 30]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002501_j.engfailanal.2020.104977-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002501_j.engfailanal.2020.104977-Figure2-1.png",
        "caption": "Fig. 2. Pneumatic tuner of torsional vibrations. 1- primary part, 2- secondary part, 3- pneumatic flexible member.",
        "texts": [
            " A variation in the inside pressure is directly linked to the variations in dynamic properties of the pneumatic tuner of torsional vibrations. The objective of this article is both theoretical and experimental investigation of heat generation and transfer in pneumatic flexible members that are used in pneumatic tuners of torsional vibrations. The design and properties of pneumatic tuners of torsional vibrations facilitate a broad array of their applications in mechanical systems. The composition of a pneumatic tuner of torsional vibrations is shown in Fig. 2. Pneumatic tuner of torsional vibrations consists of the primary (1) and secondary (2) parts which are connected by the pneumatic flexible member (3). Contrary to their linear counterparts, pneumatic flexible members used in damping of torsional vibrations are typically exposed to loading at higher frequencies. Moreover, the design of pneumatic tuners implies that pneumatic flexible members are subject to additional loading that acts multi-axially. These distinctions can affect heat generation and transfer, which leads to a temperature increase of the pneumatic flexible member"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003678_a:1008896010368-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003678_a:1008896010368-Figure9-1.png",
        "caption": "Figure 9. Stick diagram of the walking robot.",
        "texts": [
            " (40) h(q, q\u0307) =  0 \u2212(m1l2 + m2(a2 + l2)+ m3l2)l1q\u03071s1\u22122 \u2212(m1l2 + m2a2)l1q\u03071s1\u22123 \u2212m1a1l1q\u03071s1\u22124 (m1l2 + m2(a2 + l2)+ m3l2)l1q\u03072s1\u22122 0 \u2212(m1l2 + m2a2)l2q\u03073s2\u22123 \u2212m1a1l2q\u03072s2\u22124 (m1l2 + m2la2)l1q\u03073s1\u22122 m1a1l1q\u03074s1\u22124 (m1l2 + m2la2)l2q\u03073s2\u22124 m1a1l2q\u03074s2\u22124 0 m1a1l2q\u03074s3\u22124 \u2212m1a1l2q\u03073s3\u22124 0  q\u0307 +  \u2212g(m1(a1 + l1)+ 2m2l1 + m3l1)s1 \u2212g(m1l2 + m2(a2 + l2)+ m3l2)s2 \u2212g(m1l2 + m2a2)s3 \u2212gm1a1s4  . (41) (27). In this figure, the constrained force at one of the legs appears almost linear, because p\u0308r , p\u0307r , p\u0308l , and p\u0307l were ignored in controlling the forces, and hence these signals were not contained in the control inputs \u03c43 and \u03c44. Figure 9 shows the stick diagram of the support leg, traced from the video tape of the experiment. This result shows that a stable gait can be produced by controlling the trajectory of the trunk of the robot. This paper described a method for controlling a biped walking robot, based upon a formulation of constrained motion control of robots under holonomic constraints. In the formulation, the trunk of the robot was positioned with respect to the world frame. Motion of the robot was generated by providing a stable trajectory to the trunk of the robot, while controlling the constrained forces so as to prevent the feet from slipping"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002048_0142331220966427-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002048_0142331220966427-Figure6-1.png",
        "caption": "Figure 6. Descomposition of the propeller thrust.",
        "texts": [
            " Two common approaches to deal with the above challenges are performing nonlinear least-squares optimization (Johansen and Fossen, 2013) and using a nonlinear modelpredictive controller (Prach and Kayacan, 2018). However, a severe problem in the two approaches is that the computational burden imposed on the quadrotor hardware is heavy, so it takes too high costs to implement the control allocation in real time. Here, we propose a linear solution that first allocates the forces and moments to the vertical and lateral forces generated by each rotor, and then computes the tilting angles and motor speeds directly. The thrust produced by each propeller (as shown in Figure 6) can be decomposed along the vertical and lateral direction into Nv and Nl as follows Nl, i =Tis(ai)= kf n2 i s(ai) \u00f030\u00de Nv, i =Tic(ai)= kf n2 i c(ai): \u00f031\u00de The forces and moments BRWF;M are linear combinations of Nl, i and Nv, i. Therefore, equation (13) can be transformed to BRWF M =AN, \u00f032\u00de where A is a constant allocation matrix shown as (33) with dimensions of 6 3 8 and N is defined as the vector of all the vertical and lateral forces A= ffiffiffi 2 p 2 0 ffiffiffi 2 p 2 0 ffiffiffi 2 p 2 0 ffiffiffi 2 p 2 0ffiffiffi 2 p 2 0 ffiffiffi 2 p 2 0 ffiffiffi 2 p 2 0 ffiffiffi 2 p 2 0 0 1 0 1 0 1 0 1ffiffiffi 2 p 2 cd ffiffiffi 2 p 2 l ffiffiffi 2 p 2 cd ffiffiffi 2 p 2 l ffiffiffi 2 p 2 cd ffiffiffi 2 p 2 l ffiffiffi 2 p 2 cd ffiffiffi 2 p 2 l ffiffiffi 2 p 2 cd ffiffiffi 2 p 2 l ffiffiffi 2 p 2 cd ffiffi 2 p 2 l ffiffiffi 2 p 2 cd ffiffiffi 2 p 2 l ffiffiffi 2 p 2 cd ffiffiffi 2 p 2 l l cd l cd l cd l cd 2 6666666666666664 3 7777777777777775 , \u00f033\u00de N= Nl, 1,Nv, 1, "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure33-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure33-1.png",
        "caption": "Fig. 33. Rigid Type of Traction Coupling with Diaphragm Drive and Outboard Bearing",
        "texts": [
            " An ordinary outboard bearing can then be used with a flexible coupling on the driven shaft ; alternatively the outboard bearing is not required if the flexible coupling is of the self-centring type, or has a centring spigot so that the short runner shaft can act in effect like a cardan shaft and permit of considerable errors in alignment. A problem arising when the gearbox is mounted on a bell housing is that there is little room for a specially mounted outboard bearing. The \u201cdiaphragm drive\u201d shown by Fig. 33 is a convenient solution which involves mounting the coupling on a thin disk of steel bolted to the crankshaft flange in place of the conventional driving disk, a spherical centre spigot being provided to ensure concentric running, In this arrangement the coupling unit is supported at both ends in such a manner that it can act in effect as a cardan shaft and the resilience of the thin driving disk introduces a measure of flexibility to minimize the effect of crankshaft bearing wear and vibration upon the runner shaft and gearbox bearings"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002991_j.matpr.2021.05.508-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002991_j.matpr.2021.05.508-Figure2-1.png",
        "caption": "Fig. 2. Schematic of material extrusion Process [21].",
        "texts": [
            " Nowadays researchers are considering composite polymers like combinations of PLA and PC, ABS and PC, PP and PE [17,18]. In this review work the effect of process parameters on mechanical properties of ABS parts are summarized for material extrusion processes. It is widely used in 3d printing because of availability and low cost of raw materials and straightforward fabrication of complex parts [19,20]. In material extrusion process metal is selectively dispensed through a nozzle [3]. The schematic diagram of a material extrusion process is shown in Fig. 2. Material extrusion can be future classified into two categories (1) Fused filament material extrusion and (2) Pellet based extrusion process. These processes are discussed in upcoming sections. Fused filament fabrication (FFF) is a direct material extrusion method that melts filaments (thermoplastic or metal) to build a layered three-dimensional part [22]. For continuous spooling of filament minimum strain at yield should be roughly 5% [23]. The filament wire is pushed by two rollers for continuous feeding to the extrusion nozzle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000204_j.autcon.2019.102996-Figure21-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000204_j.autcon.2019.102996-Figure21-1.png",
        "caption": "Fig. 21. Instructions for assembly.",
        "texts": [
            " For this project, we tagged the elements and interconnecting parts; the remaining components were all standardized, exempting need for tags. The interconnecting parts have extruded text tags added to their geometries. The tag consists of the node identification number followed by the element identification number. The tags are meant to be oriented to the outer side of the surface. The information included in the tags for the elements is the identification number of the element, its length in millimeters, and the corresponding connecting nodes identification numbers. After the production of all components, Fig. 21 should be the only information needed to assemble the physical model. We built the proof of concept without an assembly sequence, expecting the experience to provide necessary insights for more complex models. Section 4.1 describes the experience. There are two major design phases in the development of this connection system. The first is proving the concept created in the computational environment on a small scale, and the second is developing a prototype with final materials and geometries to build a one to one scale physical model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002357_j.triboint.2020.106481-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002357_j.triboint.2020.106481-Figure4-1.png",
        "caption": "Fig. 4. Dynamic bearing model including EHL.",
        "texts": [
            " Elastic deformation For the inner ring center is radially displaced by \ud835\udeff for outer ring is assumed to be fixed, the total squeezing of a ball along the normal of the ball-ring contact center (\ud835\udc650, \ud835\udc660) can easily be calculated from Fig. 3. In this case, elastic deformation along the ball-ring contact angle direction; can be calculated as the sum of film thickness and mutual approach by Eq. (15), from the film thickness equation given in Eq. (11) [39]. (\ud835\udeff1\ud835\udc57 + \ud835\udeff 2 \ud835\udc57 ) = (\ud835\udc51\ud835\udc56,\ud835\udc57 + \ud835\udc51\ud835\udc5c,\ud835\udc57 ) \u2212 (\u210e\ud835\udc56,\ud835\udc57 + \u210e\ud835\udc5c,\ud835\udc57 ) (15) Hereby, each ball contact in the bearing can be modeled as a Kelvin\u2013Voigt element shown in Fig. 4 and the contact force can be determined iteratively for a given mutual approach by the quasi-static method [40,41]. 2.2.3. EHL Contact force The EHL contact force for the mutual approach \ud835\udeff, squeezing speed ?\u0307? the film thickness at the contact center \u210e\ud835\udc50 , the Hertzian elastic deformation at the contact center \ud835\udc51\ud835\udc50 , and the damping \ud835\udc50(\ud835\udeff)caused by Tribology International 151 (2020) 106481 H. Bal and N. Akt\u00fcrk the lubrication along the contact angle direction in each ball can be obtained by Eq. (16) [42]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000948_irc.2019.00102-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000948_irc.2019.00102-Figure2-1.png",
        "caption": "Figure 2. Modeling for the low-level push recovery controller; the ankle strategy, the hip strategy, the step strategy",
        "texts": [
            " The step strategy is balanced by changing the position of the support in the direction of perturbation. We analyzed the linearized model for each strategy. Also, we explained how it applies to humanoid robots that do not support torque control. The ankle strategy controls the actuator to maintain the CoM(Center of Mass) of the robot. Since the robot used in this paper uses actuators that do not support torque control, the formulas of torque control are approximated by the formulas for position control [7]. In the Figure 2, is the result of modeling based on ankle strategy. The linearized modeling equation is presented in (1). 488 978-1-5386-9245-5/19/$31.00 \u00a92019 IEEE DOI 10.1109/IRC.2019.00102 Authorized licensed use limited to: Fondren Library Rice University. Downloaded on May 17,2020 at 04:58:41 UTC from IEEE Xplore. Restrictions apply. (1) where is the torque value, is the mass of the LIPM, is the height of the CoM, is the horizontal distance of the CoM from the current support point, is the gravity constant",
            " (2) (3) where , , , and is the PD control gain values, is the error value of . Figure 3 shows the results of applying the system modeling equation to DARwIn-OP on the simulation where (a) shows the torso of robot being pushed, (b) shows the state after the push, and (c) shows the CoM recovery by driving the ankle actuator. The hip strategy generates the counter force for the push by using the angular acceleration of torso to pull the CoM of the robot toward the base of the support. When a human is pushed, it is similar to rotate the torso. Figure 2 shows the system modeling using a flywheel. The linearized modeling equation is (4) (5) where is the mass of the flywheel, is the height of CoM, is the rotational inertia, and is the torque applied to the center of the flywheel. However, the torso of the humanoid robot cannot rotate beyond the limit of the joint. Therefore, we use the bang-bang control technique as in [8]. (6) (7) where is the maximum torque applied to the joint, is the time at which the torso stops accelerating, is the time at which the torso stops, and is the maximum joint value",
            " After reaching the target position, the actuator corresponding to hip moves to the initial value to eliminate the oscillating error that occurs. Figure 4 shows the results of applying the system modeling equation to DARwIn-OP on the simulation where (a) shows the torso of robot being pushed, (b) shows the state after the push, and (c) shows the CoM recovery by driving the hip actuator. The step strategy can be balanced by stepping the support base in the direction of the push. It has been studied based on the concept of capture point [9], [10]. The capture point can be analyzed using the analytical solution of the flywheel model in Figure 2. (8) where is the linear velocity of torso. 489 Authorized licensed use limited to: Fondren Library Rice University. Downloaded on May 17,2020 at 04:58:41 UTC from IEEE Xplore. Restrictions apply. The equation shows that the capture point is proportional to the linear velocity of torso, and can be defined as the point at which the robot pauses to stop from the push. Figure 5 shows the results of applying the capture point to DARwIn-OP on the simulation where (a) shows the torso of robot being pushed, (b) shows the state after the push, and (c) shows the CoM recovery by stepping the support base"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000692_ecticon.2016.7561468-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000692_ecticon.2016.7561468-Figure2-1.png",
        "caption": "Fig. 2. Space vector equivalent circuit of a DFIG in the rotating reference frame.",
        "texts": [
            " For the generator operates in the sub-synchronous speed, the slip is positive, which the rotorside of the DFIG will absorb power from the utility grid through the back-to-back converter, whereas the generator operates in the super-synchronous speed, the slip is negative, the rotor power will be delivered from the rotor-side through the back-to-back converter to the utility grid. The dynamic models of the DFIG are presented considering the state of the art related to this field. Generally, the equivalent circuit of a DFIG can be expressed in the rotating reference frame, rotating in space at the synchronous speed s\u03c9 , as shown in Fig. 2. Therefore, the complete voltage and flux-linkage space vector equations of the DFIG in the rotating reference frame are represented as follows: s s s s s s r r r r sl r d v R i j dt d v R i j dt \u03bb \u03c9 \u03bb \u03bb \u03c9 \u03bb = + + = + +     (1) s s s m r r r r m s L i L i L i L i \u03bb \u03bb = + = +       (2) where ,s rv v   are the stator and rotor voltage vectors, ,s ri i   are the stator and rotor current vectors, ,s r\u03bb \u03bb   are the stator and rotor flux-linkage vectors, ,s rR R are the stator and rotor resistances, , ,s r mL L L are the stator, rotor, and magnetizing inductances, s\u03c9 is the synchronous angular speed, r\u03c9 is the rotor angular speed, and sl\u03c9 is the slip angular speed ( )sl s r\u03c9 \u03c9 \u03c9= \u2212 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003626_piae_proc_1922_017_035_02-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003626_piae_proc_1922_017_035_02-Figure11-1.png",
        "caption": "FIG. 11. F I ~ . 12.",
        "texts": [
            " and in the hands of a smptic it should be safe and efficient. The risk lies in the blind use of the aoefficients under oonditions to which they do not apply. Asaunie a wntilevm bean! of length L and breadth b and thiclknelss t with a load W at tho end. Then i t is -11 known that the deffeotion of this cantilever is If several such antilevers are superposed, it is wgued-clearly assuming no friation between them-that the deflection will be inversely as the number of aantilevers. The deflection. then. for n plates of equal size (see Fig. 11) would be Thc method of Reuleaux is briefly this. 6 = WL\u2019/3EI = 4 WL\u2019lEbt\u2019 at The University of Auckland Library on June 5, 2016pau.sagepub.comDownloaded from 494 THE INSI\u2019ITUTION OF AUTOMOBILE ENCtINEERS. Reubeaux calculated that if the plates are shortened in equal .+pc (a5 in Fig. 12) the deflection will be 50 per cent greatel, 50 that Heldt (Gasoline -4utomobile, Vol. 11. 1918), who follows Reuleaux, uses a factor to allow for the common practice of using mom than one full lfength or master leaf, and he assumes that the &- flection in an actual spring is reduced by 15 per cent on acciounlt: of fridion This is clearly orudie, and it will be appasent later that it cannd be more than a rough approximation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003770_la970030o-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003770_la970030o-Figure6-1.png",
        "caption": "Figure 6. (Top) Simplified model of the distribution of the gramicidinmonomolecular channels inasectionofphospholipid monolayer (B) sandwiched between the two conducting phases electrolyte (C)andmercuryelectrode (A), respectively.ddenotes the average distance between the pore centres, 2r is the pore diameter, and m is the pore length (from ref 8). (Bottom) Schematic model of a potential profile across a lipidmonolayer (B) adsorbed on mercury (A) to electrolyte (C) at an applied negative potential (Ea). The location of the gramicidin channel (D) in the lipid layer is shown in relation to the potential (\u03c61) on themonolayer surfaceandthepotential (\u03c62)within themouth of the channel. (a) Neutral lipid monolayer in nonadsorbing electrolyte; (b) neutral lipid monolayer in electrolyte with specific adsorbing cations; (c) mixed neutral and negatively charged lipid monolayer in nonadsorbing electrolyte.",
        "texts": [
            " Previous work also showed that DOPC layers with additive compounds but without gramicidin are totally impermeable to Tl+, showing that in every gramicidin-modified layer with additive compounds it is only thegramicidinchannelwhich conducts the Tl+.8 Cyclic voltammograms of the Tl(I)/ Tl(Hg) redox couple at gramicidin-modified pure phospholipid-coated electrodes with different concentrations of gramicidin do not show marked differences as the gramicidin concentration is increased,8 and the CV reduction current increases in a regular manner concurrent with an increase in gramicidin in the phospholipid tending to a maximum.25 This supports the contention that the gramicidin does not form significant clusters in pure DOPC layers. Figure 6 (top) summarizes as before8 thedistribution of gramicidin channels in theDOPC layer consistent with the experimental evidence. In order to understandhowmonolayer fixed charge and the nature of the lipid environment affect the results obtained in this study, it is helpful to consider the charge structure of the interface. It is assumed that the PZC of the DOPC-coated mercury electrode is similar to that of an uncoated mercury electrode, which is at -0.4 V,28 and that the average capacitance of the lipid monolayer at potentials more positive than -0",
            " This more than compensates for theelectrode chargeandproduces its ownpositive potential with respect to the electrode and solution. The incorporation of the negatively charged PS into DOPC presents a further situation in that the negative potential originating fromthePSat thePS-solution interface forms part of the total applied potential profile. These three cases are analogous to the examples of the potential profiles due to no specific adsorption, and the specific adsorption of cations and anions on a negatively charged electrode surface, respectively,36 and is summarized in Figure 6 (bottom). Models of the potential profile across a lipid monolayer are complicated and still subject to discussion. There are several contributions to the monolayer potential which originate from the fixed charge on the surface,13,14,37 the dipole potential,13,14,38-41 and the orientation of the bound water molecules.13,14 Figure 6 (bottom) considers only the potentials originating from the applied potential and the Coulombic charges at the interface as an initial approximation,assumingthat the lipidmonolayerbehaves as a low dielectric inert region and that the fixed charge is present as a layer on its surface. An applied potential falls more or less linearly along the length of the channel in the low dielectric region of the monolayer,42,43 and in the absence of a fixed charge on the monolayer surface, the major fraction of the potential drop is in this region",
            " The effect of this potential on the concentration ofTl+ in this region (cm) is givenby theBoltzmann factor:47 Within the mouth of the pore, a second potential, \u03c62, is located immediately beyond the free energybarrier of pore entry.48 This will influence the heterogeneous rate constant in the field direction for the passage of the ion across the channel-modified monolayer (kp), according to the following relation:6,48 where R is the transfer coefficient, which equals 0.5 when it operates over a symmetrical rate barrier. k0 is the heterogeneous rate constant for the passage of Tl+ across the channel-modified monolayer in the absence of an electric field. In Figure 6 (bottom) the location of these potentials is drawn with respect to the channel mouth, and it can be seen that both the fixed positive charge and the fixed negative charge will affect the values of \u03c61 and \u03c62 relative to those on DOPC in K+ electrolyte. 2. Chronoamperometry: Effect of Electrolytes, K+, Mg2+, and K+-Dy3+, and Monolayer Additives, PS, B[a]p, and Retinol. The decreased values of the intercept on the m axis compared to the theoretically predicted value of 1.52 \u00b5A s1/2 in the i versus m plots displayed in Figure 3 are interpreted as due to decreases in the monolayer surface concentration of Tl+ (cm) compared to the bulk (cTl). Table 1 displays values of the rate constant of translocation across the monolayer (kp) and values of cm/cTl. These are calculated from the i versusm plots derived from the various systems at the longer time scaleandat theappliedpotential of-0.6V. Thevariations of cm and kp can be explained by reference to the model in Figure 6 (bottom) and the influence of the potentials \u03c61 and \u03c62 on these parameters. On gramicidin-modified DOPC-coated electrodes in K+ electrolyte, the apparent decrease in concentration of Tl+ on the surface of the monolayer may correspond to a positive potential originating from the monolayer surface. This is consistent with previous reports which show that the effective ion concentration near a zwitterionic phospholipid layer (36) Stern, O. Z. Electrochem. 1924, 30, 508-516. (37) McLaughlin, S. Annu",
            " cm ) cTl exp(-\u03c61F/RT) (8) kp ) k0 exp[-R(\u03c62 - \u03c61)F/RT] (9) surface differs from that in the bulk owing to the redistribution of ions in the electrostatic, hydration, and steric-repulsion fields of the lipid layer.13 On gramicidinmodified DOPC-coated electrodes in K+ electrolyte with added Dy3+, the adsorbed Dy3+ further decreases the concentration of Tl+ on the monolayer surface as well as decreasing its translocationrate through thechannel.This can be interpreted as being caused by the increased positive potentials, \u03c61 and (\u03c62 - \u03c61) (Figure 6 (bottom) (b)). InMg2+ electrolyte, the influence onkp is not evident, since presumably the change in bulk electrolyte from K+ toMg2+ affects Tl+ transport in the channel due to factors in addition to fixed charge. On gramicidin-modified DOPC-0.4 PSmonolayers, the concentration of Tl+ on the monolayer surfaceaswell as its translocation rate through the channel is increased. This can be explained as due to the negative charge of PS,which renders the potentials \u03c61 and (\u03c62 - \u03c61) increasingly negative compared to those on gramicidin-modified DOPC monolayers in K+ electrolyte (Figure 6 (bottom) (c)). The effect of B[a]p and retinol in increasing the parameterskp and cm/cTl associatedwith the translocation of Tl+ across the gramicidin-modified layer (Table 1) can be due to a number of factors, some of which alter the potential profile across the monolayer. Firstly, these additive compounds thicken themonolayer,8,17 which can stress the monolayer-channel junction such that the monolayer surface itself forms part of the mouth of the channel which is a kind of vestibule.49,50 In this case, the applied potential will fall along this vestibule section as well as the channel itself"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001175_s0005117919090121-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001175_s0005117919090121-Figure1-1.png",
        "caption": "Fig. 1. Test self-made quadrotors: (a) robot Quadracon (FlyMaple); (b) robot Bond (TRIK).",
        "texts": [
            " Existing adaptive systems, for example [1, 20, 29], rely on Euler angles and, in general, standard PIDcontrollers. This work consists of four parts. The first defines the mathematical model that used. The second part dealing with the stabilization system and the proof of its stability. The third part provides a system for the adaptation and identification of quadrotor thrust coefficients, which are key for stable flight. The last part describes the conducted modeling of the overall system. It should also be noted that the stabilization system was tested on real, self-made quadrotors (Fig. 1), one of which worked on the FlyMap controller, the second on the TRIK controller (for more on this Russian development, see www.trikset.com, [27]). 2. MATHEMATICAL MODEL OF QUADROTOR 2.1. Designations Let q be a quaternion. Denote by qw the scalar part of q, and by qv\u2014the vector part. Then q = (qw, qv) = (qw, qx, qy, qz). Let r and s be two vectors in R 3. Their scalar product is denoted as \u3008r, s\u3009, and the vector one as r \u00d7 s. The product of quaternions a = (aw, av) and b = (bw, bv) can be written as a \u2217 b = (awbw \u2212 \u3008av, bv\u3009, awbv + bwav + av \u00d7 bv) "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000282_physreve.100.063107-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000282_physreve.100.063107-Figure2-1.png",
        "caption": "FIG. 2. Schematic of the model sperm geometry, where r represents the position vector for a point on the centerline of the flagellum. X -Y coordinates denote a fixed frame and s represents the arclength of flagellum measured from the cell body. Internal shear force field, which generates active flagellar beating, is shown by f (s, t ). Unit vectors t and n are tangent and normal vectors to the flagellar centerline, respectively. \u03c8 (s, t ) is the angle between the tangent vector and the horizontal axis.",
        "texts": [
            " THEORETICAL FRAMEWORK Using RFT, we establish the governing equations for the shape dynamics of sperm through a balance of hydrodynamic and elastic forces and internal force generation [26,52,53,55]. Many experimental observations have revealed that the flagellar waveforms of sperm from numerous organisms are approximately planar [7,13,21,49]. Thus, we consider twodimensional planar beating of a cylindrical sperm flagellum of diameter d , where the geometrical parameters of the active sperm flagellum model are detailed in Fig. 2. In the low Reynolds number approximation, the governing equation for flagellar motion was derived by equating the hydrodynamic force to the elastic force per unit length along the arclength of the flagellum as (see Fig. 2) (\u03be\u2016tt + \u03be\u22a5nn) \u00b7 { \u2202r \u2202t \u2212 U } = \u2212\u03b4G \u03b4r , (1) where the flagellum has a total arclength L, and \u03be\u2016 and \u03be\u22a5 are tangential and normal resistance coefficients per unit length of the flagellum, respectively. U is the background flow field taken here to be U = \u03b5(X \u2212 Y) with strain rate \u03b5. The total elastic energy of the flagellum G comprises contributions from bending energy, extensional energy, and active energy, where the last results from dynein motor activity [53]. Therefore, the expression for the total elastic energy of a flagellum is 063107-2 given by G = \u222b L 0 { \u03basC2 2 + \u2202r \u2202s \u00b7 \u2202r \u2202s + f } ds, (2) where \u03bas and C are the flagellar bending stiffness and local curvature, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002922_j.mechmachtheory.2021.104382-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002922_j.mechmachtheory.2021.104382-Figure1-1.png",
        "caption": "Fig. 1. 1T2R parallel manipulator: (a) CAD model and (b) kinematic scheme.",
        "texts": [
            " In Section 3 , the degree of dynamic isotropy (DDI) is defined considering pure translational DoFs, pure rotational DoFs, and mixed DoFs. In Section 4 , the effects of velocity, external load, and gravity are analyzed. The CAC index is defined as the area of the feasible acceleration region. In Section 5 , the DDI and CAC of the 1T2R parallel manipulator are analyzed, and the relationship between dynamic isotropy performance and CAC is discussed. Finally, conclusions are given in Section 6 . The computer-aided design (CAD) model and kinematic scheme of the proposed 1T2R parallel manipulator [16] are shown in Fig. 1 . The end-effector (the polishing tool) is connected to the base through three identical RCU limbs (R denotes a revolute joint, C represents a cylindrical joint, and U represents a universal joint). This parallel manipulator is actuated using three ball screw drives. The active DoFs of the manipulator are one active translational DoF (along the z -axis) and two active rotational DoFs (around the x - and y -axes). The pose of the end-effector can be expressed by (z, \u03d5, \u03b8 ) , where z is the z -coordinate of the point o \u2032 in the global coordinate system : o \u2212 xyz",
            " The parasitic motion of the end-effector is the translational motion along the x - and y -axes, which can be expressed as [16] { x = 1 2 r cos 2 \u03d5(1 \u2212 cos \u03b8 ) y = \u2212 1 2 r sin 2 \u03d5(1 \u2212 cos \u03b8 ) (6) The closed-loop vector equation is derived as shown in Eq. (7) and the inverse kinematics can be solved. o B i + B i P i = oo \u2032 + o \u2032 P i (7) To derive the velocities, accelerations of the centroid, angular velocities, and angular accelerations of the limbs, a local coordinate system is attached to each joint based on the Denavit\u2013Hartenberg method. The established coordinate systems of the first limb ( B 1 P 1 ) are shown in Fig. 1 (b). Based on the coordinate systems, the Jacobian matrix of each limb can be derived as follows. J i \u02d9 \u03be = \u02d9 x , J i = [ J v 1 , J v 2 , J v 3 , J v 4 , J v 5 ] , i = 1 , . . . , 3 (8) where \u02d9 x = [ \u02d9 x, \u02d9 y, \u02d9 z, \u03c9 x , \u03c9 y , \u03c9 z ] T is the velocity of the end-effector, \u02d9 \u03be is the velocity vector of each joint, J v j = [ z j \u00d7 ( o \u2032 \u2212 o j ) z j ] , j = 1 , 2 , 4 , 5 , and J v j = [ z j 0 3 \u00d71 ] , j = 3 . By taking the derivatives of Eq. (6) , the velocity of the parasitic motion can be expressed as { \u02d9 x = \u2212r \u02d9 \u03d5 sin 2 \u03d5(1 \u2212 cos \u03b8 ) + 1 2 r \u02d9 \u03b8 cos 2 \u03d5 sin \u03b8 \u02d9 y = \u2212r \u02d9 \u03d5 cos 2 \u03d5(1 \u2212 cos \u03b8 ) \u2212 1 2 r \u02d9 \u03b8 sin 2 \u03d5 sin \u03b8 (9) Under the description of the T&T angles, the angular velocity of the end-effector can be obtained as \u03c9 p = \u23a1 \u23a3 \u03c9 x \u03c9 y \u03c9 z \u23a4 \u23a6 = \u23a1 \u23a2 \u23a3 \u2212 \u02d9 \u03d5 cos \u03d5 sin \u03b8 \u2212 \u02d9 \u03b8sin\u03d5 \u2212 \u02d9 \u03d5 sin\u03d5 sin \u03b8 + \u02d9 \u03b8 cos \u03d5 \u02d9 \u03d5 \u2212 \u02d9 \u03d5 cos\u03b8 \u23a4 \u23a5 \u23a6 (10) By taking the derivative of Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002152_978-981-13-3549-5-Figure5.17-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002152_978-981-13-3549-5-Figure5.17-1.png",
        "caption": "Fig. 5.17 Coordinates and vectors from rover body to a wheel [101]",
        "texts": [
            " On the other hand, virtual rovers should be equipped with virtual robotic intelligence, which is developed on robotic intelligence and improved to adapt to exact virtual environment. Kinematics, dynamics and terramechanics are all essential for the high-fidelity simulation of a rover.General kinematics and dynamicsmodeling of planetary rovers are described in this subsection, and the terramechanics will be summarized in the next subsection. Coordinates and vectors from rover body to a wheel are shown in Fig. 5.17 [101]. Planetary rovers are articulatedmultibody systemswith amoving base and nw end points (wheels). Let q [q1 q2 \u00b7 \u00b7 \u00b7 qnv ]T denote the joint variables, where nv is the number of joints. Let qs [ql qm qn \u00b7 \u00b7 \u00b7 qs ]T denote a branch from the rover body to a wheel, and ns denote the number of elements in qs . Replace the joint number l, m, n, \u2026, s of the branch with 1, 2, 3 \u2026, ns. This also shows the inertial coordinates { I} and coordinates { i} attached to link i (i l, m, n, \u2026, s) and related vectors, where pi is the position vector of link i, ri is the position vector of the centroid of link i, cij is the link vector from link i to joint j, lij pj\u2212 pi is the link vector from joint i to joint j, and lie is the vector from joint i to end point e"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002781_j.matpr.2021.02.045-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002781_j.matpr.2021.02.045-Figure1-1.png",
        "caption": "Fig. 1. Solid modelling of bumper.",
        "texts": [
            " For glass fiber epoxy reinforced composites and hemp fiber reinforced composites, the material properties are evaluated. Table 1 presents the material properties of of glass fiber and hybridized glass natural fiber reinforced composites. Solid modeling is a set of specifications for numerical and PC display of three-dimensional solids and is known by its focus on real devotion from relevant regions of mathematical display and PC illustrations. By using catia, the solid model of the bumper is created and the bumper feature is applied to the model for the simulation. The Fig. 1 represents the solid modelling of bumper. The force that acts on the bumper, F \u00bc m a where, m = The vehicle\u2019s mass crashed on the bumper a = Acceleration from the crashing vehicle\u2019s gravity a \u00bc u v\u00f0 \u00de=t where, u = Until crashing, initial velocity (m\\s2) v = After crashing, final velocity (m\\s2) t = time taken a= (2.22\u20130)\\ 0.1 = 22.22 m\\s ^3 F = 1620 * 22.22 = 35996.4 N The vehicle bumper is mounted on the front of the vehicle chassis; the vehicle frame is attached to the end of the vehicle bumper"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001955_j.mechmachtheory.2020.104090-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001955_j.mechmachtheory.2020.104090-Figure1-1.png",
        "caption": "Fig. 1. Schematic of the architectures of the 6-6 SPMs discussed in this paper.",
        "texts": [
            " The main evidences in support of these conclusions are that among the analytical methods, perhaps the only method to cover all the architectures is that proposed by Husty in [25] , in 1996; and also that there are no reported examples available on two possible classes of SPMs: ones with planar FP and general MP and vice versa. This paper addresses the two issues mentioned above. It explores the geometry of the constraint surfaces of all 6-6 SPMs, and from an analysis of these, proposes a formulation which holds for any 6-6 SPM, irrespective of the planar or spatial (equivalently, general) geometry of either the FP or the MP. All the four possible combinations (shown schematically in Fig. 1 ) are discussed geometrically, and the respective FKPs are solved by reducing the constraint equations to the corresponding FKUs. To a great extent, even this reduction has been done in a uniform manner. In all the cases, the original system of six kinematic constraint equations are reduced first to a system of three polynomial equations, all of the same degree 2 in at least two of the three unknowns . Because of this special structure of the said polynomials, the remaining steps in the algorithm are again independent of the architecture",
            " Section 3 demonstrates the unified geometric formulation applicable to these manipulators, leading to the algebraic formulation of their FKPs. The methods of implementation of the algebraic formulation in a generic numerical programming language are discussed in Section 4 . The theoretical developments are illustrated in terms of SPMs of various architectures as well as that of the 6- R SS manipulator in Section 5 . The contributions of the paper are discussed in Section 6 and the conclusions are presented in Section 7 . Architectures of the 6-6 SPMs considered in this paper have been depicted schematically in Fig. 1 . As noted in Section 1 , the solutions to the FKPs of the GHSPM (shown in Fig. 1 a) and the GSPM (shown in Fig. 1 d) have been reported by many researchers. However, no solutions to the FKPs of the SPMs with the other two architectures, i.e., SPM with a non-planar MP and planar FP (see Fig. 1 b) and one with planar MP and non-planar FP (see Fig. 1 c) have been reported yet, to the best of the knowledge of the authors 5 As the solution of the FKP of the SPM with a planar MP and general FP helps solve the FKPs of other 6-6 spatial manipulators, e.g., the 6- R SS (see Fig. 2 ) it is the one chosen to be analysed in detail in this paper. Also, it turns out that the FKPs of the SPM with a planar FP and general MP is identical to that of the GSPM from an analytical perspective (see Section 3.2 ), hence it suffices to study the FKP of the latter alone"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003023_s11071-021-06591-0-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003023_s11071-021-06591-0-Figure7-1.png",
        "caption": "Fig. 7 A simplified two-dimensional plate of the head-yawing",
        "texts": [
            " Figure 6a, b show the yawing angle and yaw amplitude of the RF in long distance swimming, respectively. It can be apparently seen that the yawing angle of the RF is difficult to avoid and neglect during forward swimming, which has significant influence on the RF dynamics. Thus, the RF resistance during the forward swimming should include the frictional drag as well as the reverse thrust caused by the head-yawing. In this paper, we define the reverse thrust as obstructive thrust TO. The fish head is simplified into a two-dimensional plate without thickness, as shown in Fig. 7. Based on the momentum theorem, the increased momentum of fluid caused by the fish head-yawing should be as same as the increased RF momentum from the obstructive thrust within the sampling period Dt. Accordingly, the dynamic model can be given by: TODt \u00bc Mhvh \u00f05\u00de where Mh is the fluid mass swept by the head-yawing within Dt, and vh is the longitudinal swing speed of the fish head. Considering the traditional fish body wave function, it can be seen that the swing amplitude of the fish body becomes zero as the envelope line intersects at the origin of coordinates, which indicates that there is no oscillating at this point set as O in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001149_ab3e7a-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001149_ab3e7a-Figure5-1.png",
        "caption": "Figure 5. (a)\u2013(c) Output electric field (real part) distributions of\u0394\u03b1=0\u00b0, 15\u00b0 and 30\u00b0, respectively. (d)\u2013(f) Angular distributions of scattered far electric field intensity |E|2, corresponding to \u0394\u03b1=0\u00b0, 15\u00b0 and 30\u00b0, respectively. (g) Deflection angle \u03b8 varies with the rotation angle \u0394\u03b1 of simulation results (blue square dots) and theory prediction by equation (10) (blue solid line). The red square dots indicate corresponding efficiency \u03b7.",
        "texts": [
            " After the two metasurfaces are twisted in opposite directions with the same angle, the real component of the transmitted electric field distributions in the x-z plane, corresponding to \u0394\u03b1=0\u00b0, 15\u00b0 and 30\u00b0, are depicted in figures 5(a)\u2013(c), respectively. The deflection with various angles is obviously observed. We also calculate the far field intensity distributions, which are shown in figures 5(d)\u2013(f). It is found that the deflection angles are in good agreement with the predicted values. Finally, the linear relation between the deflection angle \u03b8 and the rotation angle \u0394\u03b1, depicted in equation (10), are verified by comparing the simulated (blue square dots) and theoretical (blue solid line) results, which is shown in figure 5(g). It is also found that the efficiency decreases with the deflection angle. Last but not least, we study the influence of the gap distance between the two metasurfaces on the deflection performance. As discussed in figure 1, d=500 nm is close enough to design the deflectors. However, it is difficult to realize for practical applications. Therefore, we calculate the far field intensity distributions for d=1 \u03bcm, 5 \u03bcm and 10 \u03bcm for the two deflectors. The deflection angle is fixed as \u03b8=15\u00b0. The results are shown in figure 6"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001019_rpj-07-2018-0182-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001019_rpj-07-2018-0182-Figure5-1.png",
        "caption": "Figure 5 Hydraulic pressure application",
        "texts": [
            " Rapid Prototyping Journal D ow nl oa de d by N ot tin gh am T re nt U ni ve rs ity A t 0 2: 47 3 1 M ay 2 01 9 (P T ) single convolution involves four steps and they are repeated till the required number of convolution is formed. Step 1: The preformed metal tube is inserted over the rubber bladder and positioned at a required height where the convolution is to be formed (Figure 3). Step 2: Top and bottom die halves are closed laterally to clamp the metal tube with the rubber bladder (Figure 4). Step 3: Application of hydraulic pressure inside the bladder expands the metal tube (Figure 5). This step is also referred as bulging (Kang et al., 2007). Step 4: The top die is closed vertically downwards till it touches the bottom die, the metal tube gets the \u201cU\u201d shape convolution (Figure 6). This step is also referred as folding (Kang et al., 2007). Then the formed convolution is shifted down and all the four steps are repeated till the required number of convolutions are achieved. Design and process parameters adapted from the previous study (Prithvirajan et al., 2017) is shown inTable I"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002471_j.engfailanal.2020.104811-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002471_j.engfailanal.2020.104811-Figure16-1.png",
        "caption": "Fig. 16. Vibration stress distribution in the CDBG with additional constraints arising from gear-teeth meshing (swing-shaped mode).",
        "texts": [
            " However, according to the natural mode analysis in Section 4.2, the vibration stress distribution in the gear disc is axisymmetric about the gear rotation axis, which is inconsistent with the crack propagation direction of the failed gear. Therefore, it is necessary to consider the low-frequency vibration stress with the additional constraint of teeth meshing. Taking into account the additional constraint ( = \u00d7k N m1 10 /c 9 ) and the 2\u00d7 rotor operational-speed-based excitation, the vibration stress distribution of the CDBG is illustrated in Fig. 16. The vibration stress is asymmetric about the gear rotation axis due to the existence of contact stiffness. Once micro cracks appear at the roots of the gear teeth, low-frequency vibration stress will cause the cracks to propagate along the dashed line in Fig. 16, which coincides with the crack shape of the failed gear in Fig. 2. In summary, the process of fatigue fracture failure of the CDBG can be described as follows: The high-frequency excitation generated by the gear-teeth meshing forces the gear system to vibrate with the 5th nodal diameter mode, causing high-cycle fatigue cracks to appear at the roots of the gear teeth. Thereafter, due to the influence of the additional constraints on the dynamics of the CDBG caused by gear-teeth meshing, the modal frequency corresponding to the swing-shaped mode increases and approaches the excitation frequency generated by the 2\u00d7 excitation arising from gas-generator rotor vibration, resulting in a low-frequency vibration stress in the gear"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001526_s00170-020-04987-7-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001526_s00170-020-04987-7-Figure2-1.png",
        "caption": "Fig. 2 Calibration artifacts a with a bidirectional scan pattern and distortion measurement locations, b with the standard rotating stripe pattern and part dimensions, and c with the rotating stripe pattern and contouring (LPBF standard scanning strategy); adapted from ANSYS\u00ae [27]",
        "texts": [
            " ASCs correspond to three direct multipliers of the inherent strain, which are applied along the local longitudinal, transverse, and depth scan directions to reflect the fact that more strain develops along the longitudinal scanning direction than in the transverse direction. A positive ASC results in compressive strains (contraction), whereas a negative ASC results in tensile strains (expansion). Note that the SSF coefficient needs to be calibrated for all strain modes, while the ASCs are only required for the scan pattern and thermal strain modes. In this study, calibration was carried out by printing three geometrically identical cross-shaped artifacts using three different scan patterns (Fig. 2). Calibration artifact no. 1 was printed using a bidirectional scan pattern along the X-axis, with a 0\u00b0 starting angle and a 0\u00b0 layer rotation angle. Calibration artifact no. 2 was printed using a rotating stripe scan pattern with a 57\u00b0 starting angle and a 67\u00b0 rotation angle. As recommended by the AP user\u2019s guide, calibration artifact nos. 1 and 2 were printed without contouring, up-skin, and down-skin. Finally, calibration artifact no. 3 was printed using the scanning strategy of a specific LPBF system (in the present work, this corresponds to the scanning strategy of an EOSINT M280/290 LPBF system). The in-plane distortions of the calibration artifacts were measured at locations where they are deemed to be maximum (Z = 22 mm, Fig. 2a). Table 1 summarizes the process parameters used for the scan pattern strain mode simulations with all three calibration artifacts. The following three-step procedure was used to calibrate the ASCs and SSF values: 1. A first series of simulations was carried out on calibration artifact no. 1. Due to its bidirectional scanning strategy, anisotropic distortions along the X- and Y-axes are expected. These unequal distortions were used to determine the longitudinal and transverse ASCs. Avalue of 1 was assumed for the depth ASC (the ASC between each layer in the Z direction) as recommended by the ANSYS Additive user guide (2018) [26]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001922_j.procir.2020.04.051-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001922_j.procir.2020.04.051-Figure3-1.png",
        "caption": "Fig. 3. Elements in FDM [19]",
        "texts": [
            " \ud835\udc64\ud835\udc64\ud835\udc64\ud835\udc64\ud835\udc56\ud835\udc56\ud835\udc56\ud835\udc56 , i=1,2,\u2026,9) are still fixed and cannot change to better fit the 9-layered part, when the relative influence in a 9-layered part is expected to be different from the relative influence in a 12-layered part. It should be noted that for the network shown in Fig. 2, the attention version is identical to the non-attention version structure-wise, with the only difference being the variable weights \ud835\udc64\ud835\udc64\ud835\udc64\ud835\udc64\ud835\udc56\ud835\udc56\ud835\udc56\ud835\udc56 , i=1,2,\u2026,N in the attention version and the fixed weights in the non-attention version. The attention mechanism-incorporated DL technique is experimentally evaluated using FDM as an example, which is illustrated in Fig. 3 [19]. The tensile strength of the printed product is chosen as the measure for part quality. The test specimen is produced according to the specification of ASTM D638-14, with the dimensions shown in Fig. 4 [20]. The FDM printer is a MakerBot Replicator with Polylactide (PLA) as the printing material [21]. Three FDM machine settings are investigated: extruder temperature (in degree C), printing speed (in mm/s) and layer height (in mm). A full factorial design is conducted for the experiment, as shown in Table 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001835_ccdc49329.2020.9164778-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001835_ccdc49329.2020.9164778-Figure1-1.png",
        "caption": "Fig 1. Geodetic coordinate system and body coordinate system The AUV dynamic model can be described as a 6 DOF\u2212 model in the earth-fixed coordinate system",
        "texts": [
            " In the second section, we give the state feedback linearization results and graph theory knowledge of the AUV model. In the third section, we give the definition of the circular formation problem. In the fourth section, we design an artificial potential field function . In Section 5, we present a distributed formation controller. In the sixth section, a simulation example is given to prove the accuracy and effectiveness of the proposed method. 2.1 State Feedback Linearization of AUV Models E \u03be\u03b7\u03b6\u2212 and the body-fixed coordinate systemO xyz\u2212 . As shown in Figure 1. Thus the nonlinear and coupled AUV model can be presented as follows: ( ) vi i iJ\u03b7 \u03b7= (v ) v (v ) v ( )i i i i i i i i i iM v C D g \u03b7 \u03c4+ + + = (1) In practical engineering, the rolling motion is generally stable and the roll amplitude is small, which can be approximated as the roll angle 0\u03c6 = and the roll angular velocity 0p = .Wherein, [ ], , , ,i i i i i ix y z\u03b7 \u03b8 \u03c8= respectively, , , , ,i i i i ix y z \u03b8 \u03c8 indicate the position of the ith AUV in the y, z direction and the Euler angle of the rotation around the axis y, z"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000049_iemdc.2019.8785171-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000049_iemdc.2019.8785171-Figure6-1.png",
        "caption": "Fig. 6. Mechanical analysis of the 18/12 switched reluctance motor (a) with the frame and (b) without the frame.",
        "texts": [],
        "surrounding_texts": [
            "The test machine is three-phase 18/12 switched reluctance motor which has the maximum power of 100 kW, the maximum torque of 209 Nm, the power density of 18 kW/L, the torque density of 35 Nm/L, the peak efficiency of 96%, with the outer stator diameter of 264 mm and the axial length of 108 mm in the active part. When the current is excited in each phase, radial and tangential forces are generated in each stator tooth. 978-1-5386-9350-6/19/$31.00 \u00a92019 IEEE 576 Figs. 1 show a part of a cross section of the 18/12 switched reluctance motor. The three stator poles are assigned as phase A, B, and C. When currents are provided, radial forces FrA, FrB, and FrC are generated as shown in Fig 1 (a). These radial forces increase as phase currents are increased. Thus, each radial force is mostly triangular in shape of square current. Radial force sum Frsum is the summation of three radial forces that is Frsum = FrA + FrB + FrC. The conventional square current generates high variation in the radial force sum causing vibration at the surface of the stator and generating high acoustic noise. The idea of flattening radial force sum is one of the solutions to reduce vibration in the stator surface. With the MinFrsum current, the radial force sum is flattened, then acoustic noise is significantly reduced. To perform flattening radial force sum, the current profile should be investigated. Fig. 1 (b) shows tangential forces in switched reluctance motor. The torque TA, TB, and TC are generated when the current in each phase is provided. The summation of all torque is the total torque Ttotal. With MinTrip current, torque ripple can be minimized. Figs. 2 (a) and (b) show the radial force and torque profiles in the 18/12 switched reluctance motor, respectively. Radial force and torque profiles are obtained by providing several dc currents from 5 A to 35 A only in one phase. The radial force and torque equations are approximated only in magnetically unsaturated condition for simplicity i.e. both the torque and radial force are assumed proportional to the square of phase current [16]. III. PROPOSED CURRENT WAVEFORMS IN THE 18/12 SRM Several current waveforms have been proposed in [16]. The most advanced current has dc, fundamental, and harmonics as, i = i0 + i1 sin(\u03b8 + \u22051) + i2 sin(2\u03b8 + \u22052) +\ud835\udc563 \ud835\udc60\ud835\udc56\ud835\udc5b(3\ud835\udf03 + \u22053) +\ud835\udc564 \ud835\udc60\ud835\udc56\ud835\udc5b(4\ud835\udf03 + \u22054) (1) The values of i0, i1, i2, i3, i4, \u03d51, \u03d52, \u03d53, and \u03d54 are gradually varied. In a look-up table, all combinations of current parameters are investigated and the notable combinations are picked up based on the criteria, such as the minimum rms current, the minimum variation of the radial force sum, and the minimum of torque ripple. Fig. 3 shows the algorithm to derive the coefficients of all the proposed currents. The numerical method was done using MATLAB. At the beginning, the Stator C FrA FrB FrC Stator Rotor C A B Frsum= FrA +FrB + FrC (a) TA TB TC Stator Rotor C A B Ttotal =TA + TB + TC FrA FrB FrC Stator Rotor C A B Frsum= FrA +FrB + FrC (b) Fig. 1. Cross sections of the 18/12 switched reluctance motor. (a) Radial forces; (b) Tangential forces. interval of current parameters is defined. There are four subroutines for torque calculation, radial force calculation, normalization, and evaluation. After all combinations of current parameters are calculated, calculated rms current, radial force sum variation, and torque ripple are listed [16]. Figs. 4 shows the notable four proposed currents with radial force sum and torque waveforms from the approximation and Maxwell analysis which have a good correspondence. The four requirements are minimum rms current (MinRMS), minimum variation of the radial force sum (MinFrsum), minimum torque ripple (MinTrip), and minimum variation of both radial force sum and torque ripple (MinBoth). These requirements are corresponding to (b), (c), (d), and (e) and colored yellow, red, green and purple in Figs. 4, respectively. The average torque is set to 5 Nm. In Figs. 4, rms current value, variation of the radial force sum \u2206Frsum, and torque ripple Trip are specified. In case of the square current, the rms current value is 10.67 Arms, variation of radial force sum \u2206Frsum = 175 N, and torque ripple Trip = 1.8 Nm corresponding with Fig. 4 (a). In MinRMS current, the minimum current is as low as 8.64 Arms, but radial force sum variation and torque ripple are significant as high as 198 N and 6.1 Nm, respectively. In MinFrsum current, radial force sum is mostly flat and rms current value is low as low as 9.91 A. In MinTrip current, torque ripple is low and rms current value is low as low as 9.4 A. In MinBoth current, both variation of the radial force sum and torque ripple are flattened and rms current value is low as low as 10.4 Arms. All criteria result in reduced rms current with respect to the square current. Fig. 5 shows frequency spectrum of radial force sum with square and the four proposed currents. Radial force sum components are indicated in blue, yellow, red, green, and purple for square, MinRMS, MinFrsum, MinTrip, and MinBoth currents, respectively. From Fig. 5, we can see that 3rd and 6th harmonic components are significantly reduced by MinFrsum and MinBoth currents. Significant of vibration and acoustic noise reductions are expected from these currents. IV. MECHANICAL ANALYSIS AND ACOUSTIC ANALYSIS After completing Maxwell analysis, the next step is performing mechanical analysis to investigate the vibration in the 18/12 switched reluctance motor. In the mechanical analysis, the result from Maxwell analysis is imported. Mechanical analysis is continuation of Maxwell analysis when investigating vibration in electrical machine. In the mechanical analysis, rotor and coil parts are suppressed, thus only stator and frame are remained. Figs. 6 (a) and (b) shows the stator body of the 18/12 switched reluctance motor with and without the frame, respectively. Vibration at the stator surface from the square and the four proposed currents is investigated. Figs. 7 shows comparison of the vibration in the stator surface between square and the four proposed currents. The 18/12 switched reluctance motor is rotated at 1000 r/min so that fundamental frequency is 200 Hz. In Figs. 7, it is shown vibration from 200 Hz to 8000 Hz for every 200 Hz steps. Vibration reduction is achieved by all the proposed currents with respect to the square current although the reduction ratio is different in each proposed current. The radial force sum variation is one of the main causes of vibration in the 18/12 switched reluctance motor thus flattened radial force sum leads to significant vibration reduction. After mechanical analysis is completed, the next step is acoustic analysis. In acoustic analysis, the area around the stator called acoustic body is investigated and all parts of the machine are suppressed. Figs. 8 (a) and (b) show acoustic body in ANSYS acoustic analysis and imported velocity at the stator surface from mechanical analysis, respectively. The hole in the center of acoustic body is suppressed 18/12 switched reluctance motor. Outer diameter of acoustic body is set to 1 meter. Figs. 9 show acoustic noise spectrum from acoustic analysis with the square and the proposed currents. Acoustic noise reduction is achieved by all the proposed currents with respect to the square current although the reduction ratio is different in each proposed current. Acoustic analysis results are corresponding with the mechanical analysis results. V. EXPERIMENTAL VERIFICATION Fig. 10 shows experiment setup configuration. From 400 V AC source, the AC voltage is rectified with AC/DC converter to 650 V DC. From the DC voltage, an inverter for load machine and the SRM inverter are connected. The 18/12 switched reluctance motor is mechanically coupled with load machine. The load machine is rated 10 kW thus only light load is possible. The acoustic noise spectrum is measured by placing microphone 30 cm in axial axis and connected with FFT analyzer. Hysteresis control is applied in the experiment system. By adjusting the band of hysteresis width, the actual Acoustic Analysis (a) Acoustic Analysis (b) Fig. 8. Acoustic analysis of the 18/12 switched reluctance motor. (a) Acoustic body; (b) Imported velocity from mechanical analysis. current waveform can follow the reference properly. In hysteresis case, switching frequency is not fixed as PWM control. Experiment was carried out to confirm effectiveness of all proposed currents. The 18/12 switched reluctance motor is rotated at 1000 r/min with average torque of 5 Nm. For each proposed current, acoustic noise spectrum is compared with the square current at the same rotational speed and average torque. Fig. 11 shows the 18/12 switched reluctance motor and the load machine. Fig. 12 shows the current waveforms of square and proposed currents in the experiment. The rms current values of square, MinRMS, MinFrsum, MinTrip, and MinBoth are 10.99 A, 9.51 A, 11.44 A, 10.9 A, and 12.81 A, respectively. The least rms current value is from MinRMS. Minimization of both variation of the radial force sum and torque ripple is also possible with the increases of rms current value by MinBoth current. However, the rms current value of MinBoth is as high as 12.81 A. Fig. 13 shows acoustic noise spectrum comparison between the square and all of the proposed currents in the experiment. From Fig. 13, we can see that the four proposed currents can reduce acoustic noise in the high frequency after 2000 Hz. It is possible because in the proposed currents, there is no sudden change in current. For the acoustic noise in low frequency range, the effect of radial force is significant. MinRMS current indicated in yellow has high acoustic noise at 3rd and 6th harmonic components, similar with the square current indicated in blue because variation of radial force sum using square and MinRMS currents is high. However, compared with MinFrsum current, significant acoustic noise reduction is achieved in 3rd and 6th harmonic components because variation of the radial force sum is minimized. Moreover, significant reduction of 3rd component is observed with MinBoth current. The 3rd harmonic component is almost gone. From these results, it is shown that the four proposed currents can reduce the acoustic noise compared with square current. The low frequency components 600 Hz in MinFrsum and MinBoth are also reduced with respect to MinRMS. Acoustic analysis results shown in Fig. 9 and acoustic noise spectrum from experiments shown in Fig. 13 have a good correspondence. The significant peaks are observed in multiple of third harmonic, such as 3rd, 6th, 9th, etc. Acoustic noise reduction with proposed currents also show a good correspondence between acoustic analysis and experiment. For example, let us see the acoustic noise between square and MinRMS currents at third harmonic component which is 600 Hz. The acoustic noise from acoustic analysis between square and MinRMS currents are almost similar. The same result is also observed in the experiment. At 600 Hz, square and MinRMS currents generate almost similar acoustic noise which is around 50 dB. Another example is acoustic noise comparison between square and MinFrsum currents. The acoustic noise from acoustic analysis at 600 Hz is reduced 20 dB by MinFrsum current with respect to the square current. Moreover, acoustic noise at 600 Hz from the experiment is also reduced 10 dB with MinFrsum current. Acoustic noise reduction is observed, although the reduction scale is different between acoustic analysis and the experiment. The difference is caused by different current. In acoustic noise analysis, ideal current is provided. However, in the experiment, current waveform is provided from hysteresis regulation. Moreover, the four proposed currents can reduce acoustic noise spectrum especially in higher frequency compared with square current. From these results, it can be seen that acoustic analysis can predict acoustic noise spectrum in the electrical machine accurately. VI. CONCLUSION In this paper, electro-mechanical analysis, acoustic analysis, and experiment are carried out to verify acoustic noise and vibration reduction with novel current waveforms. The novel current waveforms include up to 4th harmonic components with phase shift in all harmonic components. The 18/12 switched reluctance motor is rotated at 1000 r/min with average torque of 5 Nm. From mechanical analysis, acoustic analysis, and experiment results, vibration and acoustic noise reductions are confirmed by using novel current waveforms with respect to the square current. Reasonable acoustic noise reduction is achieved with MinFrsum and MinBoth currents because of small variation in the radial force sum."
        ]
    },
    {
        "image_filename": "designv11_14_0002646_s0263574720001290-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002646_s0263574720001290-Figure3-1.png",
        "caption": "Fig. 3. Pitch motion desired pose of 3-R(RRR)R+R HAM.",
        "texts": [
            " Altogether, k, \u03c2 and r are the pitch angle, azimuth angle and relative fixed coordinate distance of the dynamic coordinate system of the antenna, respectively. The angle \u03b86 between Tz-axis and Cx -axis represents the polarization angle. In the process of antenna tracking satellite, pitching and azimuth motion are usually required. The pitching and azimuth motion process of the 3-R(RRR)R+R HAM is analyzed. During the pitching motion, the azimuth of the antenna remains unchanged. The initial pose of the 3-R(RRR)R+R HAM is the initial pose for antenna movement, as shown in Fig. 2. The expected pose of the pitch motion is shown in Fig. 3. In the initial pose, a certain vector T V is given in the end coordinate system Tx T Ty and the direction of the vector is along the intersection line of plane OCT and plane Tx T Ty . The direction of vector T V is always along the intersection of plane OCT and plane Tx T Ty , and the rotation angle of the polarization mechanism is \u03b86 = 0\u25e6. Therefore, in the process of pitching, not only the antenna reflector does not move with it but also the polarization mechanism remains stationary. In the course of azimuth motion, the antenna pitch angle remains unchanged. As shown in Fig. 3, the expected pose of pitch motion is selected as the initial pose of azimuth motion and Fig. 4 is set as the desired pose of azimuth motion. In the course of azimuth motion, the HAM changes from the initial pose to the expected pose and the azimuth changes from \u03c2 to \u03c21. To keep the direction of vector T V always along the intersection of plane OCT and plane Tx T Ty , the rotation angle of the polarization mechanism is \u03b86. Two sets of transition poses can be used to realize the expected pose of the HAM from the initial pose of the azimuth motion to the expected pose, as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001096_s00170-019-04076-4-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001096_s00170-019-04076-4-Figure3-1.png",
        "caption": "Fig. 3 Coordinate system O-xyz representing relationship between axes of grinding wheel and worm",
        "texts": [
            " 4 Relationship between axes of worm and worm wheel where es \u00bc d f 1 2 \u00fe H\u2212 dm1 2 This curved line is rotated by angle \u03c8 about zg axis. Therefore, the coordinates of the point P on the curved surface of the grinding wheel in Og-xgygzg is expressed as a position vector: Xg u;\u03c8\u00f0 \u00de \u00bc \u2212 u\u2212Rm\u00f0 \u00desin\u03c8 u\u2212Rm\u00f0 \u00decos\u03c8ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u03c12\u2212 es\u2212u\u00f0 \u00de2 q \u2212 V 2 1 2 6664 3 7775 \u00f02\u00de The curved surface of the grinding wheel is defined by a combination (u, \u03c8). The rectangularcoordinate system O-xyz is set based on the relationship between the axes of the grinding wheel and worm as shown in Fig. 3. The worm axis coincides with z axis. The axis zg of grinding wheel is inclined by \u03b3 about y axis. The common perpendicular of the axes of the grinding wheel and worm is y axis. The center distance between the axes of the grinding wheel and worm is e. The point P on the curved surface in Og-xgygzg in Fig. 2 is transformed into the coordinate system O-xyz. That is, the point P in Og-xgygzg is moved along yg axis by e, it is rotated about yg axis by angle \u03b3. Therefore, the coordinates of point P in O-xyz after this procedure is expressed as a position vector: X u;\u03c8\u00f0 \u00de \u00bc MX g u;\u03c8\u00f0 \u00de \u00f03\u00de whereM is the 4 \u00d7 4 matrix of the rotational and translational coordinate transformation from Og-xgygzg to O-xyz and is represented by M \u00bc cos\u03b3 0 0 1 sin\u03b3 0 0 e \u2212sin\u03b3 0 0 0 cos\u03b3 0 0 1 2 64 3 75 \u00f04\u00de Assuming the relative velocity W between the grinding wheel and worm, and the unit surface normal N of X, the equation of meshing between the grinding wheel and worm is as follows [17, 18]: N u;\u03c8\u00f0 \u00de\u2219W u;\u03c8\u00f0 \u00de \u00bc 0 \u00f05\u00de N(u, \u03c8) and W(u, \u03c8) in Eq",
            " (5) are determined by N u;\u03c8\u00f0 \u00de \u00bc \u2202X \u2202u \u2202X \u2202\u03c8 \u2202X \u2202u \u2202X \u2202\u03c8 W u;\u03c8\u00f0 \u00de \u00bc \u03c9 k X \u00fe hk\u00f0 \u00de where k is the unit vector toward z axis and \u03c9 is the relative angular velocity between the not rotating grinding wheel and rotating worm. Moreover, h is the screw parameter of the worm. From Eq. (5), we can obtain \u03c8 corresponding to the depth of cut u. X can be determined when these values are substituted for the right side in Eq. (3). Assuming that x, y, and z components of X are defined as A, B, and C, respectively, these mean the generating line of the worm tooth surface. Moreover, when \u03b8 is the rotation angle of the screw motion about the worm axis in Fig. 3, the coordinates of the right tooth surface of worm is expressed as a position vector: X r u;\u03c8; \u03b8\u00f0 \u00de \u00bc xr yr zr 1 2 64 3 75 \u00bc Acos\u03b8\u2212Bsin\u03b8 Asin\u03b8\u00fe Bcos\u03b8 h \u03b8\u2212\u03b80\u00f0 \u00de \u00fe C 1 2 64 3 75 \u00f06\u00de where \u03b80 is the initial value of \u03b8. The unit surface normal ofXr is represented by Nr. Figure 4 shows the relationship between the axes of the worm and worm wheel whose angular velocities are \u03c91 and \u03c92, respectively. The common perpendicular of the axes of the worm and worm wheel is y axis, too. The offset distance between the worm and worm wheel is defined as E"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001096_s00170-019-04076-4-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001096_s00170-019-04076-4-Figure7-1.png",
        "caption": "Fig. 7 5-axis machining center",
        "texts": [
            " Then, the tooth surface deviation of the worm wheel is measured and the worm wheel may be machined again according to the sizes of the deviations although this case hardly occurs, too. Finally, the machined worm and worm wheel are meshed each other and the experimental tooth contact pattern is observed in order to confirm the validity of this machining method. The worm tooth surface is machined using an offset machining method. In this case, CAD/CAM system is not required and NC data is created using the analyzed results of contact point between the worm tooth surface and the side surface of the end mill. A 5-axis machining center (Okuma B750) was used as shown in Fig. 7 [22]. This machine consists of 3 axes (x, y, and z axes) of straight lines, and the axis of the tool inclination (B axis) and the rotation axis of workpiece (C axis), namely, total 5 axes. However, only two z and C axis control of the 5-axis control are used in machining of worm. In this case, the worm is machined producing screw motion with two z and C axis control. Figure 8 shows the method of offset machining of the right tooth surface of the worm in xz cross section. The end mill and worm tooth surface contact at point Q and the coordinates of point Q in O-xyz is expressed as a position vector Xr in Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001209_jfm.2019.681-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001209_jfm.2019.681-Figure1-1.png",
        "caption": "FIGURE 1. (Colour online) (a) Schematic diagram of a multilayered hollow sphere submerged in an infinite non-Newtonian fluid accounting for both shear and compressional relaxation effects. (b) Geometric schematic of an anisotropic N-ply hollow sphere.",
        "texts": [
            " A cc es s pa id b y th e U CS F Li br ar y, o n 05 O ct 2 01 9 at 1 3: 11 :4 6, s ub je ct to th e Ca m br id ge C or e te rm s of u se , a va ila bl e at h tt ps :// w w w .c am br id ge .o rg /c or e/ te rm s. that additional damping mechanisms intrinsic to the solid particles also should be examined. This paper develops a comprehensive theoretical analysis of the 3-D coupled vibration of a multilayered anisotropic hollow sphere submerged in an infinite non-Newtonian fluid characterized by a compressional linear viscoelastic model proposed by Yong (2014) and Chakraborty & Sader (2015), as illustrated in figure 1. Full FSI is taken into consideration by imposing the continuity conditions at the fluid\u2013solid interface. Numerical examples are finally conducted to investigate the effects of different factors, including fluid viscosity and compressibility, fluid viscoelasticity, solid anisotropy, solid surface effect and solid intrinsic damping, on the vibration characteristics of the FSI system. The study can be envisioned as a generalization of the existing works mentioned above. The obtained 3-D analytical solution covers a few simpler and degenerate cases",
            " However, the compressible linear Maxwell model only considers the shear relaxation effect and overlooks the compressional relaxation process that may play a significant role in the dynamic response of a pressure wave in a non-Newtonian fluid. As a mathematical and physical modification to the linear Maxwell model, Yong (2014) and Chakraborty & Sader (2015) proposed a compressible linear viscoelastic model accounting for both shear and compressional relaxation effects, which is characterized by a spring\u2013dashpot connected in series for each of the shear and compressional relaxation processes, as displayed in figure 1(a). The compressible linear viscoelastic model will be referred to as the compressional non-Newtonian model hereafter for simplicity, which can recover a purely Newtonian result in the low-frequency limit and reproduce a purely elastic response in the high-frequency limit (Chakraborty & Sader 2015). In this paper, we assume that the fluid is non-Newtonian, and can be described by the compressional non-Newtonian model. Furthermore, the linearized governing equations of the compressible viscoelastic fluid are solved by introducing appropriate potential functions in the case of small-amplitude vibrations",
            " A cc es s pa id b y th e U CS F Li br ar y, o n 05 O ct 2 01 9 at 1 3: 11 :4 6, s ub je ct to th e Ca m br id ge C or e te rm s of u se , a va ila bl e at h tt ps :// w w w .c am br id ge .o rg /c or e/ te rm s. Substituting (3.4) into (3.1)\u2013(3.3), through some lengthy mathematical manipulations, one can transform the original 3-D equations into two separated state equations, both with variable coefficients. The reader is referred to the supplementary material SM-I available at https://doi.org/10.1017/jfm.2019.681 for the derivations of the state equations for the general non-axisymmetric free vibration of a closed N-ply hollow sphere shown in figure 1(b). Since it is intractable to solve the state equations with variable coefficients directly, the approximate laminate technique (Chen & Ding 2002; Ding, Chen & Zhang 2006; Wu et al. 2017) can be employed to derive the approximate analytical solutions and then obtain the relation between the state vectors at the inner and outer surfaces of the N-ply hollow sphere (see SM-I for more information of the approximate laminate technique). As a result, two separate transfer relations can be established as Tin kn = HknTou kn, k \u2208 {1, 2}, n > 1, (3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001915_s41872-020-00151-y-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001915_s41872-020-00151-y-Figure1-1.png",
        "caption": "Fig. 1 Schematic of the experimental rig",
        "texts": [
            " The orientation of the paper is in this manner: Section\u00a02 describes the dataset used, proposed methodology, and feature extraction process. Section\u00a03 briefly describes methods used for classification. A brief description of performance parameters is given in Sect.\u00a04. Section\u00a05 discusses the results, and finally, the conclusion is given in Sect.\u00a06. The vibration signals for the study are acquired from the Case Western Reserve University (CWRU) bearing data center (Loparo 2019). The schematic view of the experimental setup is shown in Fig.\u00a01. Three classes of signals are obtained for various faults namely: (a) Inner race fault (b) ball fault (c) outer race fault at 1730, 1750, 1772, and 1797\u00a0rpm of the motor. The fault diameter considered for different classes are 0.007, 0.014 and 0.021 inches, respectively, and depth to be 0.011 inches. Time responses of 1 3 signals are shown in Fig.\u00a02 and the reconstructed signal for level 7 decomposition is given in Fig.\u00a03. Through discretization of the wavelet \u2217 (a,b) (t) , DWT is derived from continuous wavelet transform (CWT)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002103_s11431-020-1718-4-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002103_s11431-020-1718-4-Figure5-1.png",
        "caption": "Figure 5 (Color online) Mechanism of electro-oxidation of glucose at hierarchical nanoporous CuO modified GCE.",
        "texts": [
            " The concentration of the interfering species is fixed 1/10 of the glucose despite their actual lower ratio (1/40 to 1/30) to glucose in human blood. As shown in Figure 4(f), a significant and steady current increase is detected while 1 mM glucose is added to the solution. Interfering signals are only 1.7% for AA and 5.9% for UA in comparison with glucose, indicating strong anti- interference capability of present glucose sensor. The mechanism for improving glucose electro-oxidation by optimizing micro-environment is illustrated in Figure 5. With the presence of glucose in alkaline solution, previous oxidized Cu (III) is served as active oxidant to react with adsorbed glucose to form gluconolactone and is reduced to Cu (II), resulting in Jpa at 0.55 V. Generally, the electrooxidation of glucose is highly depended on the nanostructure of CuO. By comparing the apparent constant of reaction (ks) of CuO with different morphology (Figure S3), the kinetics of glucose oxidation process is highly enhanced due to hieratical nanostructure. On the one hand, a great number of firecracker-shaped CuO nanorods provide high sensitivity by increasing the surface area (Figure S2) and the number of active sites for glucose\u2019s electro-oxidation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001509_icicict46008.2019.8993271-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001509_icicict46008.2019.8993271-Figure4-1.png",
        "caption": "Fig. 4. A triangle and corresponding local features.",
        "texts": [
            " Features Extraction We denote the ith triangle of a Delaunay triangulation net by Ti = {ma, mb, mc}, mk |k\u2208{a,b,c} = {xk, yk, \u03b8k, tk}, where minutiae ma, mb, mc are vertexes of the triangle, (xk, yk) is the coordinates of the minutia mk, \u03b8k is the orientation of its associated edge, and tk \u2208 {0, 1} is the minutia type (0 corresponds to ridge ending while 1 corresponds to ridge bifurcation). Unlike [41], we do not use minutia type in our scheme due to its instability, so the feature vector of Ti is expressed by FVi = {dab, dbc, dca, \u03b1ab, \u03b1bc, \u03b1ca} dab = \u221a (xa \u2212 xb)2 + (ya \u2212 yb)2 dbc = \u221a (xb \u2212 xc)2 + (yb \u2212 yc)2 dca = \u221a (xc \u2212 xa)2 + (yc \u2212 ya)2 \u03b1ab = tan\u22121 ( ya \u2212 yb xa \u2212 xb ) \u2212 \u03b8a \u03b1bc = tan\u22121 ( yb \u2212 yc xb \u2212 xc ) \u2212 \u03b8b \u03b1ca = tan\u22121 ( yc \u2212 ya xc \u2212 xa ) \u2212 \u03b8c The triangle Ti and its features are demonstrated in Figure 4. Suppose there are s triangles in the Delaunay triangulation net, then the fingerprint image can be expressed by a set of these s local feature vectors as SV = {FVi }si=1. In our construction, to reduce matching processing time, we choose the first 80 Delaunay triangles (s = 80) from the whole set in ascending order of the distance between them and the singular point or the center of the fingerprint image. Both d and \u03b1 are quantized and represented as bit strings of length 4, so FVi can be represented by a 24-bit binary string"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002721_j.aej.2021.01.012-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002721_j.aej.2021.01.012-Figure2-1.png",
        "caption": "Fig. 2 Fragment of different profiles of a worm thread in the axial cross-section.",
        "texts": [
            " Mxy is a homogeneous transformation matrix from the system y to x, whereas M 0 xy is a homogeneous transformation matrix connected with the rotation of the model in the direction from the system y to x. For x and y, an appropriate index related to the coordinate system number is inserted. 3. Mathematical model of the hourglass worm hob The description of the flank surface of machining worm thread is based on defining an analytical equation of a globoid helix and a parametric equation of the tool profile. It is assumed that the axial profile of the worm hob is rectilinear or arcshaped (concave or convex) (Fig. 2). The profile is limited by points B y1B; z1B\u00f0 \u00de and C\u00f0y1C; z1C\u00de. These are the coordinates of the tooth profile vertices on one tooth flank. The parametric equation of the BC section has the following form: x1\u00f0u\u00de \u00bc 0 y1\u00f0u\u00de \u00bc y1B \u00fe \u00f0y1C y1B\u00de u z1\u00f0u\u00de \u00bc z1B \u00fe \u00f0z1C z1B\u00de u \u00f02\u00de where u is the equation parameter (up u uk, up \u00bc 0, uk \u00bc 1). In the case of the arc profile, the equation of any given arc in the plane of y1z1 is represented as x1\u00f0h\u00de \u00bc 0 y1 h\u00f0 \u00de \u00bc R cos h\u00f0 \u00de \u00fe y0 z1 h\u00f0 \u00de \u00bc R sin h\u00f0 \u00de \u00fe z0 \u00f03\u00de where h is the equation parameter (hp h hk\u00de, y0and z0 are the coordinates of the arc profile centre, and R is the arc radius",
            " The coordinates of the profile limiting points B andC can be obtained after assuming and determining the basic geometric parameters of the globoid worm gearing based on the AGMA or GOST standard and the equations presented in this paper [19]. In the case of the arc profile, the arc radius R must be assumed. The coordinates y0 and z0 have to be determined by using the parametric equation of the circle passing through two points, which in this case are points B andC. The equation parameters hp and hkcan be obtained based on trigonometric relationships as shown in Fig. 2. The parametric equation of the tooth flank surface of the worm hob is obtained by developing a tooth profile along the globoid helix (Fig. 3). The position vector of the tooth flank surface of worm hob with the rectilinear axial profile is determined by r 1 0\u00f0 \u00de 1 \u00bc M 1 0 1 M12 M2 0 2 M21 x1 u\u00f0 \u00de y1 u\u00f0 \u00de z1 u\u00f0 \u00de 1 2 6664 3 7775 \u00f04\u00de After development, one gets the following form: r 1 0\u00f0 \u00de 1 u1; u\u00f0 \u00de \u00bc x1\u00f0u\u00de cos u1\u00f0 \u00de a sin u1\u00f0 \u00de \u00fe a cos u2\u00f0 \u00de sin u1\u00f0 \u00de\u00fe \u00fey1\u00f0u\u00de cos u2\u00f0 \u00de sin u1\u00f0 \u00de z1\u00f0u\u00de sin u2\u00f0 \u00de sin u1\u00f0 \u00de x1\u00f0u\u00de sin u1\u00f0 \u00de a cos u1\u00f0 \u00de \u00fe a cos u1\u00f0 \u00de cos u2\u00f0 \u00de\u00fe \u00fey1\u00f0u\u00de cos u2\u00f0 \u00de cos u1\u00f0 \u00de z1\u00f0u\u00de sin u2\u00f0 \u00de cos u1\u00f0 \u00de a sin u2\u00f0 \u00de \u00fe y1\u00f0u\u00de sin u2\u00f0 \u00de \u00fe z1\u00f0u\u00de cos u2\u00f0 \u00de 1 2 6666666666666666666664 3 7777777777777777777775 \u00f05\u00de In the case of the tool with arc axial profile, the position vector of the surface assumes the following form: r 1 0\u00f0 \u00de 1 \u00bc M 1 0 1 M12 M2 0 2 M21 x1 h\u00f0 \u00de y1 h\u00f0 \u00de z1 h\u00f0 \u00de 1 2 6664 3 7775 \u00f06\u00de After development, one gets the following equation: r 1 0\u00f0 \u00de 1 u1; u\u00f0 \u00de \u00bc x1\u00f0h\u00de cos u1\u00f0 \u00de a sin u1\u00f0 \u00de \u00fe a cos u2\u00f0 \u00de sin u1\u00f0 \u00de\u00fe \u00fey1\u00f0h\u00de cos u2\u00f0 \u00de sin u1\u00f0 \u00de z1\u00f0h\u00de sin u2\u00f0 \u00de sin u1\u00f0 \u00de x1\u00f0h\u00de sin u1\u00f0 \u00de a cos u1\u00f0 \u00de \u00fe a cos u1\u00f0 \u00de cos u2\u00f0 \u00de\u00fe \u00fey1\u00f0h\u00de cos u2\u00f0 \u00de cos u1\u00f0 \u00de z1\u00f0h\u00de sin u2\u00f0 \u00de cos u1\u00f0 \u00de a sin u2\u00f0 \u00de \u00fe y1\u00f0h\u00de sin u2\u00f0 \u00de \u00fe z1\u00f0h\u00de cos u2\u00f0 \u00de 1 2 6666666666666664 3 7777777777777775 \u00f07\u00de The parameter u1 in the above equations denotes the worm hob thread length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001207_rpj-07-2018-0171-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001207_rpj-07-2018-0171-Figure6-1.png",
        "caption": "Figure 6 (a) Model for manufacturing with allowances in STEP, (b) smoothed model in STL, (c) model with supports",
        "texts": [
            " The final mesh contained approximately 205,000 elements and 386,000 nodes, whereas 80 per cent of the elements had element quality better than 0.5. In terms of degrees of freedom (DOFs), the model size was approximately 1.15 106 with threeDOFs per node. A direct solver, time step of 0.1 s, unsymmetrical Newton\u2013Rhampson method, constant stabilization, weak springs and 5 per cent convergence on the total deformation of the axle carrier were set for the analysis. The final CAD model [Figure 3(d)] was added with machining allowances of minimum 1mm [Figure 6(a)] and exported in the STL format. The polygonal mesh was then finally smoothed in the Autodesk Meshmixer (Autodesk, Inc., San Rafael, CA, USA) to eliminate remaining sharp edges and minimize the stress concentrators caused by defects of the mesh [Figure 6(b)]. Data for additive manufacturing were prepared in software Materialise Magics 21.11 (Materialise, Leuven, Belgium). The axle carrier was placed on a building platform with nearly minimal height to avoid residual stress and further deformation in the central area used for mounting the bearings. Support structures of two types were used [Figure 6(c)]. Block supports for overhanging geometries and solid supports for preventing the component from deflection during the build job. It also enhances heat dissipation during component building. The axle carrier prototype was fabricated using the SLM280HL machine (SLM Solutions Group AG, L\u00fcbeck, Germany) with standard process parameters: laser power of 350W, scanning speed of 930mm/s, hatch distance of 0.17mm and layer thickness of 50mm. The build job consisted of 1,575 layers, whereas the total build time was 17h"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002077_j.mechmachtheory.2020.104209-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002077_j.mechmachtheory.2020.104209-Figure2-1.png",
        "caption": "Fig. 2. Kinetostatic model of an Exechon-like PKM.",
        "texts": [
            " With the above Cartesian coordinate settings, the transformation matrixes between the body-fixed frames P-uvw, A i -x i y i z i and the global reference frame O-xyz can be derived as O P : T rans ( P \u2212 u v w \u2192 O \u2212 xyz ) (1) O Ai : Trans ( A i \u2212 x i y i z i \u2192 O \u2212 xyz ) (2) where O P and O Ai denote the transformation matrixes of P-uvw and A i -x i y i z i with respect to O-xyz , respectively, which can be referred to our previous publication [40] . The Exechon-like PKM is simplified into an equivalent kinetostatic model, in which all passive joints are treated as virtual spring units with lumped-parameter method while the limb body is modeled as an elastic spatial beam through the finite element method [41] . The Exechon-like PKM shown in Fig. 1 (a) can be simplified into the equivalent kinetostatic model as depicted in Fig. 2 . As depicted in Fig. 2 , the kinetostatic model of the Exechon-like PKM is assumed with the modular components that mainly consist of the elastic joints and limbs as well as the rigid platform and base. The passive joints are represented by a set of virtual spring units with equivalent translational/torsional stiffness constants of k L Ai / k A Ai and k L Bi / k A Bi , whose corre- sponding linear/angular elastic deformations are denoted as x L Ai / x A Ai and x L Bi / x A Bi , respectively. The equivalent spring constants of passive joints of the Exechon-like PKM can be further defined in a general form as k L Ai = diag [ k L Aix k L Aiy k L Aiz ] , k A Ai = diag [ k A Aix k A Aiy k A Aiz ] (3) k L Bi = diag [ k L Bix k L Biy k L Biz ] , k A Bi = diag [ k A Bix k A Biy k A Biz ] (4) where k L Aix / k A Aix , k L Aiy / k A Aiy and k L Aiz / k A Aiz denote the linear/angular stiffness coefficients of the passive joint at A i along/about corresponding coordinate axis in the frame A i -x i y i z i ",
            " Thus, when the prismatic joint and the motor are locked the linear stiffness coefficients of the B i joint along z i axis can be rewritten as k L Biz 0 = [ ( k L Biz )\u22121 + ( d i \u2212 l ni ) / EA + k \u22121 Fbi + k \u22121 Ri ] \u22121 ; k Ri = [ k \u22121 Rbi + ( l d / 2 \u03c0) 2 \u00b7 k \u22121 Moi ]\u22121 (5) where d i represents the equivalent distance between P joint and A i joint; l ni ( i = 1, 2, 3) represents the length from front bearing to A i joint; \u2018E\u2019, \u2018A\u2019 and l d are the Young\u2019s modulus, cross-sectional area and lead of the lead screw, respectively; k Fbi and k Rbi denote the axial stiffness coefficients of the front bearing and the rear bearing in the lead screw assemblage in the i th limb, respectively; k Moi denotes the equivalent torsional stiffness coefficient of the locked motor about the lead screw. These stiffness coefficients of bearings and motors can be determined through its specifications and datasheets. In Fig. 2 , the l P 0 and l i 0 represent the equivalent vectors from the gravity center of the platform and the i th ( i = 1, 2, 3) limb assemblage to the centre P of the platform, respectively. The external force and moment F P and M P (including of inherent gravities) lead to the elastic deformations of the PKM. The equivalent gravities of the platform and the i th ( i = 1, 2, 3) limb assemblage imposed on their gravity centers can be expressed in the frame O-xyz as G P = m p0 g , G Li = m li g ( i = 1 , 2 , 3 ) (6) where m p 0 denotes the mass of the moving platform; m li denotes the mass of the i th ( i = 1, 2, 3) limb assemblage; g is the gravitational acceleration",
            " In addition, F M denotes the equivalent external load acting on the moving platform and can be expressed as F M = [ F T P M T P ]T (24) The Exechon-like PKM is designed for high speed machining where the external cutting force acting on the platform is usually much smaller than its own gravity. In such a case, the gravity contributes the most to the kinetostatic behaviors [40] . Therefore, the external cutting force is neglected and only the gravity is included in the established kinetostatic analysis. The equivalent external load as described in Eq. (24) turns into the follow F P = G P + 3 \u2211 i =1 G Li ; M P = l P0 \u00d7 G P + 3 \u2211 i =1 l i 0 \u00d7 G Li (25) where l P 0 and l i 0 denote the position vectors of gravity centers as illustrated in Fig. 2 . By substituting Eq. (25) into Eq. (18) , the elastic displacement X N \u00d71 of the PKM as described in Eq. (19) can be solved. With Eq. (19) , the deflection X M of the moving platform can be obtained as the last 6 \u00d7 1 column vector in X N \u00d71 . Hence, the gravity-caused deflections of the platform of the Exechon-like PKM can be given by ( X M ) 6 \u00d71 = X ( N \u22125: N ) = [ \u03b5 Px \u03b5 Py \u03b5 Pz \u03bePx \u03bePy \u03bePz ]T (26) where X ( N -5: N ) denotes the 6 \u00d7 1 column vector of the X N \u00d71 from N- 5 to N ; \u03b5 Pk and \u03bePk ( k = x, y, z ) denote the linear and the angular displacements of the platform along and about the k th Cartesian coordinate axis, respectively"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003517_tf9524800526-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003517_tf9524800526-Figure1-1.png",
        "caption": "FIG. 1 .-Quenching of chlorophyll a fluorescence",
        "texts": [
            " Aliquot portions were injected from a syringe into a more dilute chlorophyll solution than finally desired and excess solvent evaporated off to a calibration mark to promote rapid degassing. Chlorophyll was prepared from spinach by a modification 4 of the method of Comar and Zscheile.4 of 25-30\" C. RESULTS POLAR soLvmTs.-On addition of phenylhydrazine, the fluorescence of chlorophyll a, in methanol was quenched according to the SternVolmer equation, i.e. IfllfO = 1/(1 + W l ) , in methanol by phenylhydrazine. Concentration of chlorophyll a, 5 x 10-6 mole 1.-1. as shown in fig. 1 by the linearity of the reciprocal of relative fluorescence intensity against concentration. The quenching constant 3.2 moles-1 1. is in substantial agreement with the value of 3.7 moles-1 1. previously reported 6 from less numerous measurements. Quenching required several minutes after addition of phenylhydrazine for the attainment of the constant quenched intensity. This time-lag serves as a differentiation between a chemical dark reaction of unexcited dye and quencher and a second-order collision deactivation of excited dye by quencher"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001234_s00202-019-00857-y-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001234_s00202-019-00857-y-Figure2-1.png",
        "caption": "Fig. 2 Connection of the supercapacitor in the DC/DC converter circuit",
        "texts": [
            "\u00a040 into 39, the rotor voltage becomes: So, the voltage in the rotor rotational reference frame is: The rotor transient resistance becomes: The stator resistance and the rotor transient resistance in Eq.\u00a043 are directly proportional. Thus, the stator flux damping is expedited, as the rotor inrush currents are decreased. Implementing a supercapacitor to a DC bus, the grid-side converter works as an active power source. Connecting ESS to a DC bus can be done either directly or via an interface. Use of a supercapacitor modeling in DFIG is given in Fig.\u00a02. Here, the connection was done via a 2-quadrant DC/ DC converter. As ESS sets generator output power with (37) i \u2217 dr2 + i \u2217 drn + i \u2217 drf = \u2212 Ls RsLm ( vds2 + ws( ds2 + dsn + dqsf) \u2212 Rs Ls ( ds2 + dsn + dsf) ) (38) i \u2217 qr2 + i \u2217 qrn + i \u2217 qrf = \u2212 Ls RsLm ( vqs2 \u2212 ws( qs2 + qsn + qsf) \u2212 Rs Ls ( qs2 + qsn + qsf) ) . (39)vdqr = Rridqr + Lr didqr dt + Lm Ls d dt dqs (40) dqs dt = vdqs \u2212 jwr dqs \u2212 Rs dqs Ls + RsLmidqr Ls (41) vdqr = ( Rr + L2 m L2 s Rs ) idqr + Lr didqr dt + Lm Ls \u00d7 ( vdqs \u2212 jws \u2212 Rs dqs Ls ) (42) vdqr = R \ufffd r idqr + Lr didqr dt + Lm Ls \u00d7 ( vdqs \u2212 jws \u2212 Rs dqs Ls ) (43)R \ufffd r = Rr + L2 m L2 s Rs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure20-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure20-1.png",
        "caption": "Fig. 20 Cutter B after grinding",
        "texts": [
            " Based on the above settings, the grinding path is generated by the method of isoparametric-line [20] as shown in Fig. 16. This experiment uses Walter five-axis tool grinding machine to grind the cutter faces. The post-processing is done according to the structure and parameters of this machine to generate applicable NC program. Then, the cutter faces are ground as shown in Fig. 17 and Fig. 18. After grinding, the cutter with curved rake face (cutter A) is obtained as shown in Fig. 19. The cutter with plane rake face (cutter B) is also manufactured for comparison as shown in Fig. 20. Then, the manufacturing accuracy is measured by the measuring module of the tool grinding machine as shown in Fig. 21 and a cutter detector as shown in Fig. 22. The radial runout is 0.0045 mm, and the end runout is 0.0047 mm. The edge shape error is 0.0045 mm. These results can fully meet the accuracy requirement. In order to verify the correctness of the cutter design and manufacture methods proposed in this paper, a cutting experiment is carried out as shown in Fig. 23. The experiment is divided into two groups corresponding to cutter A or cutter B"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001096_s00170-019-04076-4-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001096_s00170-019-04076-4-Figure8-1.png",
        "caption": "Fig. 8 Offset machining of worm",
        "texts": [
            " In this case, CAD/CAM system is not required and NC data is created using the analyzed results of contact point between the worm tooth surface and the side surface of the end mill. A 5-axis machining center (Okuma B750) was used as shown in Fig. 7 [22]. This machine consists of 3 axes (x, y, and z axes) of straight lines, and the axis of the tool inclination (B axis) and the rotation axis of workpiece (C axis), namely, total 5 axes. However, only two z and C axis control of the 5-axis control are used in machining of worm. In this case, the worm is machined producing screw motion with two z and C axis control. Figure 8 shows the method of offset machining of the right tooth surface of the worm in xz cross section. The end mill and worm tooth surface contact at point Q and the coordinates of point Q in O-xyz is expressed as a position vector Xr in Eq. (6). The offset distance T is defined as the distance between the worm axis and the axis of the end mill. The unit surface normal Nr of the tooth surface Xr passes through the axis of the center of the end mill because of the swarf cutting. y component ny of Nr is always zero as shown in Fig. 8 using only two z and C axis control. In this case, \u03b8 is obtained when u is provided because ny is functions of u and \u03b8. Since xr in Eq. (6) is x component of the right tooth surface of the worm, the following equation which represents the right side of tooth space yields: T \u00bc Acos\u03b8\u2212Bsin\u03b8\u00fe d 2 nx \u00f09\u00de where d is the diameter of the end mill and nx is the x component of Nr. A and B in Eq. (9) are the function of u, and nx is the functions of u and \u03b8. Since \u03b8 is obtained when u is provided from ny = 0, T is determined substituting the obtained u and \u03b8 for Eq"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002541_0954407020974497-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002541_0954407020974497-Figure3-1.png",
        "caption": "Figure 3. Representation of steering linkage as two slider crank chains.",
        "texts": [
            " For Ackerman less than 100%, the IC will lie beyond the wheel base towards the rear axis. As mentioned above, for the vehicle to negotiate a smooth cornering, the angles turned by inside and outside wheel in the rack and pinion steering geometry should satisfy Ackerman condition. The rack and pinion type steering geometry consists of two tie rods, two tie arms and one steering rack. This steering linkage geometry resembles the combination of two identical slider crank mechanisms. The right hand side mechanism is the mirror image of the left hand side mechanism as presented in Figure 3 drawn as a schematic top-view representation of the vehicle\u2019s steering system. Tie-arm, tie-rod and rack denote the six moving links of two slider crank chains viz. two cranks, two connecting rods and two sliders. The translational motion of the rack simulates the linear motion of two sliders and constrained to move along the same axis. a and b are the lengths of crank and connecting rod respectively. Let xi and xo are the instantaneous displacements of inner and outer sliders corresponding to the tie arm\u2019s input fi and fo"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002329_042033-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002329_042033-Figure3-1.png",
        "caption": "Figure 3. Virtual stand.",
        "texts": [
            " To simulate the lateral displacement of the tractor resulting from tire deformation, the wheel model is also supplemented with elastic elements (figure 2, b). ICMSIT 2020 Journal of Physics: Conference Series 1515 (2020) 042033 IOP Publishing doi:10.1088/1742-6596/1515/4/042033 A virtual spring 4 connects the wheel disk and the tire having the possibility of axial displacement without relative rotation. By selecting the elastic coefficients and damping of the virtual spring, tire deformation can be simulated. To study the lateral stability, the machine-tractor unit is installed on a virtual stand (figure 3). It consists of a fixed base and a platform that changes the angle of inclination. The interaction between the contact pairs \u201cplatform \u2013 wheels\u201d had the parameters of the standard SolidWorks Motion interaction \u201csteel \u2013 rubber\u201d. During the simulation, the following parameters were monitored: \u2013 the angle of inclination of the platform; \u2013 the angle of the tractor; \u2013 the contact force of the front and rear wheels distant from the axis of inclination of the platform (figure 4). To check the adequacy of the tractor simulation model, a virtual experiment was conducted to study the transverse static stability of the MTZ-82 tractor (figure 5, b)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001761_0954405420928684-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001761_0954405420928684-Figure3-1.png",
        "caption": "Figure 3. CAD file and selective laser melting (SLM) printing of the specimen used in (a) preliminary and (b) further experiments.",
        "texts": [
            " While the addition of the filler materials is anticipated to alter the mechanical properties of the laserconsolidated parts,40 experiments are designed to establish the role of the Si3N4 filler fibres both on the laser consolidation mechanics of 316L and the ensuing mechanical properties. Preliminary work was done to establish the working ranges of the SLM process parameters, mainly laser power and speed. In order to avoid the compounding effects, the laser scan speed was initially kept constant at 750 mm/s, while the laser power was varied from 175 W to 225 W, in steps of 25 W. All these preliminary tests were conducted based on cubic samples of 103 103 10 mm size. Figure 3(a) presents the placement of the cubic samples on the build plate. The scan strategy involved zig-zag motion of the laser energy source from one end to the centre of the sample, covering the section area in two parallel patches. The scan lines and the sections used for the metallography studies are indicated in the inset. It may be pertinent to point out here that this kind of scan strategy leaves a poorly coalesced zone along the centre line of each section because of the heating of adjacent powders in differential times",
            " Based on the results of these initial trials, more elaborate experimental plans were developed for subsequent experiments, printing specimens for both metallographic and mechanical characterisation. Considering the time-consuming and expensive printing involved, a multi-specimen CAD model was designed to simultaneously print three tensile bars and six metallography samples for each process parameter combination. This printing strategy allowed for three repetitions of each tensile test and microstructural evaluation in a particular direction to be undertaken for statistical validation of the experimental results. Eight of these models were stacked parallel to one another as depicted in Figure 3(b), so that 83 3=24 sets of samples could be printed in each build, thus saving considerably on the pre- and postprinting tasks as well as the actual printing time. Printing for the whole experimental plan was completed in just two builds, one with 316L and the other with 316L plus Si3N4. Selective laser melting was done based on the Renishaw system at the Auckland University of Technology, New Zealand (Renishaw AM 400). The AM 400 system is equipped with a solid-state Nd:YAG laser (wavelength=1070 nm) of 75 mm spot size",
            " The austenite formation followed cellular dendritic growth in the build direction, and the orientation of the cells changed in adjacent fusion zones, as the laser scan direction changed from one track to the other. The low magnification optical microscopy results obtained from three sets of each of 316L and 316LSi3N4 laser-melted samples are presented one below the other at different laser power settings in Figure 6. The images are on the cross sections of the cubic samples at right angles to the centre line of separation between the two raster patches on each section, as depicted in Figure 3(a). The layer-wise consolidation of the metal powders was evident from the microstructures. The lack-of-fusion problem is evidently central to the section in the cases corresponding to the last two columns of the first-row images. This is apparent from the scan strategy followed, which uses zig-zag raster paths that switch directions central to the specimen, as depicted in the inset in Figure 3(a). Also, a lack of interlayer fusion and a varying degree of delamination are evident at higher power ranges used with the 316L powder material. The presence of the Si3N4 nanoparticles appears to have overcome this problem to a large extent, as the microstructures in all the images of the bottom row of Figure 6 show a homogenised uniform dispersion of flat layers oriented parallel to the base plate and with strong inter-layer fusion and structural integration. Also, the favourable role of the presence of the nano- ceramic particles appears to increase with increasing laser power"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002352_j.mechmachtheory.2020.103945-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002352_j.mechmachtheory.2020.103945-Figure5-1.png",
        "caption": "Fig. 5. \u201c(5R)\u201d modules: (a) \u201c \u02dc R u 1 \u22a5 \u2194 R u 2 \u22a5 \u0303 R u 3 \u201d, (b) \u201c \u2194 R u 1 \u22a5 \u0303 R u 2 / \u0303 R u 3 \u201d and (c) \u201c \u02dc R u 1 / \u0303 R u 2  \u0303 R u 3 \u201d.",
        "texts": [],
        "surrounding_texts": [
            "W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx 3\nwhere M sl and n sl denote respectively the DOF of the external single-loop linkage and the number of R pairs of the linkage, and C cc and n cc denote respectively the number of the constraints contributed by the coupling chain and the number of R pairs of the chain. From Eq. (1) , n sl and n cc can be enumerated in detail, listed in Table 1 .\nIn the recent work [34 , 35] , two classes of deployable units with n sl = 8 and n cc = 5, and n sl = 9 and n cc = 4 have been synthesized. The remaining three classes of units with n sl = 10 and n cc = 3, n sl = 11 and n cc = 2, and n sl = 12 and n cc = 1 will be developed here.\nA two-layer and two-loop deployable linkage unit needs to have a plane-symmetrical structure, so its R pairs need to be arranged symmetrically with respect to the symmetrical plane (SP). As a result, for the deployable unit with n sl = 10 and n cc = 3, there exist only two kinds of structure types, denoted as \u201c{4R} [3R] (6R)\u201d and \u201c{6R} [3R] (4R)\u201d, shown in Fig. 1 (b) and (c), in which FB denotes the fixed base, and MP1 and MP2 denote the two links connecting the coupling chain with the external loop linkage. A \u201c{4R} [3R] (6R)\u201d unit can be decomposed into a 4R lower module (denoted as \u201c{4R}\u201d), a 3R coupling chain module (denoted as \u201c[3R]\u201d), and a 6R upper module (denoted as \u201c(6R)\u201d). Similarly, a \u201c{6R} [3R] (4R)\u201d\nunit can be decomposed into a lower module \u201c{6R}\u201d, a coupling chain module \u201c[3R]\u201d and an upper module \u201c(4R)\u201d.\nFor the deployable unit with n sl = 11 and n cc = 2, there exists only one structure type shown in Fig. 1 (d), denoted as\n\u201c{6R} [2R] (5R)\u201d. Such a unit can be decomposed into three modules \u201c{6R}\u201d, \u201c[2R]\u201d and \u201c(5R)\u201d.\nFor the deployable unit with n sl = 12 and n cc = 1, there is only one structure type shown in Fig. 1 (e), denoted as \u201c{6R} [1R]\n(6R)\u201d. Such a unit can be decomposed into three modules \u201c{6R}\u201d, \u201c[1R]\u201d and \u201c(6R)\u201d.\n2.2. Structure modules of new deployable linkage units\nThe above linkage units involve two kinds of lower modules\u201c{4R}\u201d and \u201c{6R}\u201d, three kinds of coupling chain modules\u201c[3R]\u201d, \u201c[2R]\u201d and \u201c[1R]\u201d, and three kinds of upper modules\u201c(6R)\u201d,\u201c(5R)\u201d and \u201c(4R)\u201d. Synthesis of those modules is the important basis of developing the three classes of linkage units. For convenience, in what follows, \u201c \u2194\nR i \u201d denotes that R i is\nlaid along the centerline of a link connected by the pair, \u201c \u02dc R i \u201d denotes that R i is perpendicular to the centerline of a link connected by the pair. \u201cR i \u201d denotes that R i is in the SP. In addition, notations \u201c/\u201d, \u201c\u22a5 \u201d and \u201c \u201ddenote the parallel, perpendicular and noncoplanar relations between joint axes, respectively.\nBy considering plane-symmetry and enumerating different joint layouts, topological structures of each kind of module can be synthesized, and the related results are presented as follows, in each of which the layouts of only R pairs in the left\nof the SP and in the SP.\n2.2.1. Coupling chain modules\nThere are three \u201c[3R]\u201d modules denoted as \u201c \u02dc R c1 / \u0303 R c2 \u201d, \u201c \u02dc R c1  \u0303 R c2 \u201d and \u201c \u2194 R c1 \u22a5 \u0303 R c2 \u201d, shown in Fig. 2 (a\u2013c). Besides, there are only one \u201c[2R]\u201d module denoted as \u201c \u02dc R c1 \u201d and only one \u201c[1R]\u201d module denoted as \u201c \u02dc R c1 \u201d, shown in Fig. 2 (d) and (e), respectively.\n2.2.2. Lower modules\nThere are two kinds of \u201c{4R}\u201d modules, \u201c \u02dc R l1  \u0303 R l2 \u201d and \u201c \u02dc R l1 \u22a5\n\u2194 R l2 \u201d, shown in Fig. 3 (a) and (b), and there are three kinds of\n\u201c{6R}\u201d modules, \u201c \u02dc R l1 \u22a5\n\u2194 R l2 \u22a5 \u0303 R l3 \u201d, \u201c \u02dc R l1  \u0303 R l2 \u22a5 \u2194 R l3 \u201d and \u201c \u02dc R l1  \u0303 R l2 / \u0303 R l3 \u201d, shown in Fig. 3 (c\u2013e), respectively.\nPlease cite this article as: W.-a. Cao, Z. Jing and H. Ding, A general method for kinematics analysis of two-layer and twoloop deployable linkages with coupling chains, Mechanism and Machine Theory, https://doi.org/10.1016/j.mechmachtheory. 2020.103945",
            "4 W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx\nRu1 MP1 MP2\nRu2 Ru4\nRu3 Ru1 MP1 MP2\nRu2 Ru4\nRu3 Ru1\nMP1 MP2\nRu2 Ru4\nRu3\nRu2\nRu1 Ru3\nRu4\nMP1 MP2\n(a) (b) (c) (d)",
            "W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx 5\nFurther, some units should be excluded due to their structural defects, such as local DOFs and undesired angle ranges[35]. Finally, there are fifteen available \u201c{4R} [3R] (6R)\u201d deployable units, listed in Table 2 , in which two typical units are shown in Fig. 7 (a) and (b).\nSimilarly, seventeen available \u201c{6R} [3R] (4R)\u201ddeployable units can be obtained, listed in Table 3 , in which two typical\nunits are shown in Fig. 7 (c) and (d).\nBased on the combination of three kinds of \u201c{6R}\u201d modules, one kind of \u201c[2R]\u201d module and three kinds of \u201c(5R)\u201d modules, there are nine \u201c{6R} [3R] (5R)\u201d units in total. After excluding those units with structural defects, six\navailable\u201c{6R} [2R] (5R)\u201d units can be obtained, listed in Table 4 , in which two typical units are shown in Fig. 8 (a) and (b). Similarly, there are also six available \u201c{6R} [1R] (6R)\u201ddeployable units, listed in Table 5 , in which two typical units are shown in Fig. 8 (c) and (d).\nPlease cite this article as: W.-a. Cao, Z. Jing and H. Ding, A general method for kinematics analysis of two-layer and twoloop deployable linkages with coupling chains, Mechanism and Machine Theory, https://doi.org/10.1016/j.mechmachtheory. 2020.103945"
        ]
    },
    {
        "image_filename": "designv11_14_0003900_jsvi.1997.1323-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003900_jsvi.1997.1323-Figure5-1.png",
        "caption": "Figure 5. Graphical interpretation of the successive substitution method.",
        "texts": [
            " The steady state equilibrium position of the system can be found by solving the following static equations, K \u00b7 r=H(V, r, 0)+Q+C0 \u2212K \u00b7 ar . (14) These equations are non-linear and implicit, and can only be solved by numerical methods. The equations are not a polynomial type. Therefore, a successive substitution method is suitable to solve it numerically [16]. The recursive procedure is defined by K \u00b7 rj+1 =H(V, rj , 0)+Q+C0 \u2212K \u00b7 ar . (15) The successive iterations are interpreted graphically in Figure 5 (r* refers to the solution of equation (14)). Convergence will certainly occur in the case of Figure 5(a), while in the case of Figure 5(b), the iterations of equation (15) will fail to converge to solution r*. Equation (15) can be modified to make the substitutions a successful process [17]. The modification which is made here is as follows, K \u00b7 r+Pm \u00b7 K \u00b7 r=Pm \u00b7 K \u00b7 r+H(V, r, 0)+Q+C0 \u2212K \u00b7 ar , (16) where Pm is a coefficient to be chosen to make the iterations convergent. The authors have found that by choosing a proper value of Pm , the iterations can always converge to an equilibrium position. After an equilibrium position is found, the linearization of the equations of motion is achieved by linearizing the hydrodynamic forces in a vicinity of the equilibrium position by perturbation methods. The displacement vector and the hydrodynamic forces can be written in the following forms in the vicinity of an equilibrium position, r= r0 +Dr, H+H0 +DH, (17) where r0 refers to the equilibrium position (r* in Figure 5), and H0 the hydrodynamic forces when the rotor system is at the equilibrium position. DH is assumed to be linearly proportional to the displacements and velocities of the system in the vicinity of the equilibrium position, that is, DH=\u2212KH \u00b7 Dr\u2212DH \u00b7 Dr\u0307, (18) where KH and DH are the matrices of the coefficients, which can be defined as stiffness and damping matrices of the oil films. Introducing equation (17) and equation (18) into equation (12), and taking the condition of the equilibrium position (14) into account, we have the linearized equations of motion of the rotor-bearing system, M \u00b7 Dr\u0308+DH \u00b7 Dr\u0307+(K+KH ) \u00b7 Dr=F+DC"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000978_iemdc.2019.8785203-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000978_iemdc.2019.8785203-Figure1-1.png",
        "caption": "Fig. 1. FBG strcture and sensing principal",
        "texts": [
            " It is found that the proposed method enables recognition of both thermal and mechanical operating conditions relevant to diagnostic purposes that are sensed by a single FBG head, thus providing a step forward in understanding the application potential of FBG technology in providing reduced cost yet improved robustness and fidelity alternative to conventional bearing monitoring techniques in electric machinery. II. OPERATING PRINCIPLES OF FIBRE BRAGG GRATING SENSOR An FBG sensor is a micro structure imprinted into the core of a standard single-mode optical fibre; a single imprint is referred to as an FBG head. It is formed longitudinally on the optical fibre core in a length typically of the order of a few millimetres, in which a modulated periodic refractive index is formed in a fibre core when exposed to an interference pattern of ultraviolet laser light [1]. Fig. 1 illustrates the structure and the fundamental operating concept of an FBG sensor. The fibre containing an FBG head is illuminated with broadband light and a particular wavelength (that meets Bragg condition) reflected by the FBG head. The FBG reflects a specific light spectrum that matches its designed Bragg wavelength. The process of fibre light excitation and examination of the reflected spectrum for Bragg wavelengths is managed by an interrogator unit [17]. The basic operation principle of FBG sensing is to monitor the narrowband reflected Bragg wavelength after injecting broadband light into the optical fibre"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002095_s11837-020-04491-z-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002095_s11837-020-04491-z-Figure1-1.png",
        "caption": "Fig. 1. Schematic of DXR setup for L-PBF experiments at 32-ID of the APS.",
        "texts": [
            "25 have demonstrated the setup for in situ monitoring the miniature LPBF process, which consists of a ytterbium fiber laser (IPG YLR-500-AC; IPG Photonics, Oxford, USA, wavelength of 1070 nm, maximum output power of 520 W) placed directly above the sample, a 17-4 PH stainless steel thin-slab sample (500 lm thickness 9 2.89 mm height) sandwiched by two glassy carbon plates (1 mm thickness 9 3 mm height), and a 17-4 PH stainless steel powder layer (approx. 100 lm height) placed on top of the base plate (Fig. 1). Note that the thickness refers to the dimension along the direction of the x-ray beam and the height refers to the dimension along the direction of laser penetration. All the laser melting experiments were conducted inside an argon (1 atm)-filled chamber to replicate as far as possible the conditions of a powder-added single-bead experiment in a L-PBF machine. The miniature powder bed sample was positioned in the path of the x-ray beam and a high-speed imaging camera placed 310 mm downstream from the sample (Fig. 1). This study used a laser spot size of 67.8 lm, a frame rate of 50,000 frames per second, and polychromatic x-rays with the first harmonic energy of 24.4 keV and wavelength centered on 0.508 A\u030a. DXR videos were processed frame by frame in ImageJ41 to optimize the brightness and to enhance the contrast to further distinguish the regions of interest, e.g., melt pool, vapor cavity, pores, etc. Figure 2 and Table I show the volume-weighted powder size distributions of the four 17-4 PH stainless steel powders from the lXCT measurements",
            " Although the lXCT porosity results show a strong correlation between the populations of entrapped gas in the powders and spherical porosity in the asbuilt parts, the porosity formation mechanism is more clearly established if the actual porosity transfer process is directly observed. Thanks to the high temporal and spatial resolution offered by DXR, the evolution of in-part spherical porosity can be traced back to entrapped gas in the powder. Figure 7a\u2013d shows an example of porosity transfer that occurred within a 500-ls time span. As demonstrated in Fig. 1, Fig. 7 shows the view along the xray direction with the laser moving from left to right indicated by the scan direction. The two traced particles fell into the melt pool and transferred their entrapped gas pores into the base plate 460 ls after the laser scanned over them: this moment is labeled as t0 (Fig. 7a). One pore was entrapped by the solidification front while the other escaped with a much longer path before leaving the liquid near the vapor cavity. The two different porosity evolutions match the pore elimination mechanism proposed by Mohammad et al"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003566_pime_proc_1945_153_010_02-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003566_pime_proc_1945_153_010_02-Figure16-1.png",
        "caption": "Fig. 16. Flange, Tread, and Rail Contacts for Various Tyre Profiles",
        "texts": [
            " at CAMBRIDGE UNIV LIBRARY on June 4, 2016pme.sagepub.comDownloaded from 38 COMMUNICATIONS ON R I D I N G AND WEARING QUALITIES O F RAILWAY CARRIAGE TYRES The 1 in 100 coned tyre profiles .used in the tests were a simple development of the present standard 1 in 20 coned profile, derived by joining a tread, tapered 1 in 100, tangentially to the root of the standard flange. Some consequences of running that modification on a rail canted at 1 in 20 to the vertical were shown, for new rail and tyre profiles, by Fig. 16 on which the standard 1 in 20 coned profile also appeared for comparison. The 1 in 100 coned profile, riding on the outer part of the rail, was lifted higher than the 1 in 20 coned profile, with the results, not only that the rail made contact farther down the flange (as mentioned in the paper) but also that the tip of the flange projected a somewhat shorter distance below rail level, and that the clearance between the rail and the flange when the wheels were central was somewhat increased. The present standard flange (contour A, British Standard Specification No. 276,1927) presumably had the optimum depth and shape respectively to guard against derailment and allow the correct lateral play of a pair of wheels within the rail gauge. The flange associated with 1 in 100 coned tread should, therefore, extend to the same radial distance below rail level, and should allow the same minimum transverse play. Those features could be achieved, as indicated in Fig. 16, by a flange profile of the shape ABCD whereby, in effect, the present standard 1 in 20 coned flange in correct relation to the rail was combined with the 1 in 100 coned tread, the only new portion of profile being the short length AB in the throat of the flange which, incidentally, did not make contact with the rail so long as the profiles were unworn. That suggested flange for 1 in 100 coned tyres extended about inch farther below rail level than the simple modified profile used in the tests. The common tangent where it made contact with the rounded corner of the rail was inclined at about 50 deg"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002329_042033-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002329_042033-Figure1-1.png",
        "caption": "Figure 1. Simulation model of the MTZ-82.1 tractor: (a) \u2013 detailed model; (b) \u2013 simplified model.",
        "texts": [
            "ADAMS, SolidWorks Motion, LMS Virtual.lab, SimPack, TruckSim) are used. They allow one to study static stability taking into account the dynamics of the object's behavior [5-15]. ICMSIT 2020 Journal of Physics: Conference Series 1515 (2020) 042033 IOP Publishing doi:10.1088/1742-6596/1515/4/042033 In the presented study, to create a tractor simulation model, the virtual modeling method was used in CAD SolidWorks and the SolidWorks Motion program. At the first stage, a simplified model of the MTZ-82.1 tractor was developed (figure 1). All stationary elements of the tractor were excluded from the model (shown in the figure in wireframe). To replace them, we used an equal mass ball of custom high density material. By adjusting the position of the ball, the center of mass of the simplified model is combined with the real operational center of mass of the tractor. Further, for the layout of the simulation model of the tractor, simplified models of the front and rear three-point linkage have been created. They have a simple parameterized geometry and retain mass-inertial parameters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003907_nme.1620380611-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003907_nme.1620380611-Figure3-1.png",
        "caption": "Figure 3. Ring elastically supported at radial direction",
        "texts": [
            " The programs take into consideration (1) the effects of the axial and shear deformations, (2) presence of an elastic and continuous support and (3) any arbitrary loading. As a special case, making the pitch angle zero, the planar systems can also be analysed by the same program. All the example problems considered here are chosen from the relevant literature, especially those which revealed some sort of difficulty with so-far available formulations. Example I: Ring elastically supported at radial direction (See Figure 3). The results obtained from the model of this study for the ring (or for the long cylindrical pipe) under the single load, P, in its plane are shown in Table I. These results are generally in close agreement with those of the literature.6 The author believes that the numerical values given in the literature for U, and M b for the cases: 4 = NO\", 4 = 45\" and 4 = 315\", respectively, which are rewritten here in Table I, are incorrect probably due to misprinting. Example ZI: Helicoidal staircase with constant cross-section"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000814_icarcv.2016.7838833-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000814_icarcv.2016.7838833-Figure2-1.png",
        "caption": "Fig. 2. Path-following errors of a nonholonomic autonomous forklift",
        "texts": [
            " Using spline curves for approximation of the path length [16], for each path parameter \u03c8 , xd(\u03c8) and yd(\u03c8) can be calculated. Assuming the desired velocity of path following, vd , and path parameter, \u03c8 as the distance travelled along the path, the following dynamics is amended to the kinematic model (3): \u03c8k+1 = \u03c8k+ vdk , |vdk |< vdmax (12) where vdk is an extra input to be determined by the controller at time step k. It should be noted that based on the above equation, forklift can traverse the path in reverse direction. As shown in Fig. 2, the path-following error is defined similar to [10] as the normal deviation from the desired path, \u03b5c. However, the calculation of the reference path parameter \u03c8r and therefore the error \u03b5c is not practical in real-time implementations. As a result, \u03c8k which is defined in (12) is used as an approximation of \u03c8r. With this approximation, the path-following errors are obtained as: \u03b5\u0302c = (xk\u2212 xd(\u03c8k))sin\u03c6(\u03c8k)\u2212 (yk\u2212 yd(\u03c8k))cos\u03c6(\u03c8k), (13) \u03b5\u0302 l = (xd(\u03c8k)\u2212 xk)cos\u03c6(\u03c8k)\u2212 (yk\u2212 yd(\u03c8k))sin\u03c6(\u03c8k), (14) \u03b5\u0302o = \u03c6(\u03c8k)\u2212\u03b8 , (15) and \u03c6(\u03c8k) = arctan ( yd(\u03c8k)\u2212 yd(\u03c8k\u22121) xd(\u03c8k)\u2212 xd(\u03c8k\u22121) ) , (16) where \u03b5\u0302c, \u03b5\u0302 l and \u03b5\u0302o are approximations of the cross-track, longitudinal and orientational errors"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001607_0954407020909663-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001607_0954407020909663-Figure5-1.png",
        "caption": "Figure 5. The hard points after optimization.",
        "texts": [],
        "surrounding_texts": [
            "Comparison of optimization variables and degree of understeer before and after optimization Comparison of optimization variables before and after optimization. Figure 6 shows that the height of the roll center of the front suspension after optimization can decline from the original 96.75 to 86.65mm. The roll steer coefficient is \u2202d=\u2202Fr. As can be seen from curves 1 and 2 in Figure 7, the coefficient before optimization is 0.131 and after optimization is 0.122. The roll steering coefficient decreased by 0.009, with little change. Comparison of equivalent cornering stiffness Cp before and after optimization. Table 9 presents the geometric suspension changes caused by the optimization of hard point and bushing stiffness have a direct influence on the equivalent cornering stiffness Cp of the suspension. Equivalent Cp of the front suspension changes from 82.69% to 75.63% after optimization, thereby falling into the reasonable range. Comparison of the degree of understeer before and after optimization. Table 10 shows that the tire cornering steer, roll steer, and lateral force steer are reduced to a certain degree via modification of the coordinates of the suspension hard points and bushing stiffness. Moreover, the degree of vehicle understeer is reduced, thus realizing the optimization of the excessive degree of vehicle understeer. Comparative analysis of vehicle minimum time handling and stability before and after optimization A simulation test on double-lane change is conducted using the ADAMS software. Based on the driver busy degree, rollover risk, side-slip risk indexes, and running and adhesion properties, single and comprehensive evaluation indexes of minimum time handling and stability are proposed in the previous literature.17,18 The formula is as follows ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s Figure 8 presents that the evaluation index of driver burden is reduced after optimization, and the degrees of driver busy and heaviness are mitigated. Figure 9 shows that after optimization, the evaluation index of rollover hazard is reduced. The roll angle is reduced Table 5. Sensitivity analysis of hard points. Hard point Roll center height (effect %) Roll steer coefficient (effect %) Hpl_tierod_inner x 1.97 2.14 Hpl_tierod_inner y 4.67 84.86 Hpl_tierod_inner z \u201396.33 \u201318.05 Hpl_tierod_outer x \u20138.43 13.68 Hpl_tierod_outer y \u20131.28 27.43 Hpl_tierod_outer z 91.2 \u201358.76 Hpl_lca_outer x 13.58 \u201329.78 Hpl_lca_outer y 4.63 \u201337.63 Hpl_lca_outer z \u201382.72 \u201342.56 Hpl_lca_front x 0.81 21.59 Hpl_lca_front y \u20133.86 6.47 Hpl_lca_front z 37.42 \u201335.42 Hpl_lca_rear x \u20131.66 22.14 Hpl_lca_rear y \u20136.38 17.85 Hpl_lca_rear z 58.07 33.83 Hpl_strut_lwr_mount x \u20132.96 \u201316.43 Hpl_strut_lwr_mount y 11.83 7.98 Hpl_strut_lwr_mount z 0.02 \u201370.32 Hpl_wheel_center x \u20130.53 52.60 Hpl_wheel_center y 5.73 \u20132.86 Hpl_wheel_center z 0.96 68.22 and rollover stability are improved. Figure 10 illustrates that the evaluation index of vehicle side-slip hazard is reduced. Hence, side-slip stability is further enhanced. Figure 11 shows that the evaluation index of dynamic is reduced. Vehicle speed and acceleration performance are improved, and vehicle ability on stable dynamic takeoff is reinforced. Figure 12 presents that the evaluation index of tire road holding is reduced. The tire camber angle is reduced, the adhesive rate is elevated, and the tire ground-grabbing ability is improved. Figure 13 illustrates that the responsiveness evaluation index is reduced, thereby improving vehicle acceleration response performance. Figure 14 shows that the comprehensive evaluation index of vehicle minimum time handling is reduced. Hence, minimum time handling and stability is improved."
        ]
    },
    {
        "image_filename": "designv11_14_0000204_j.autcon.2019.102996-Figure19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000204_j.autcon.2019.102996-Figure19-1.png",
        "caption": "Fig. 19. Center point (CP) is modified to show influence on local geometry of the interconnecting part.",
        "texts": [
            " The radius of the internal cylinder is the same as the radius of the hub, and the radius of the external cylinder is larger than the internal one by any amount necessary for the transition. The pipe created at the defined intersection point extends through a vector of any amplitude desired, in the direction of the element centerline; it is the extension vector. Mesh vertices provide the connection center points; mesh normal vectors (hub axis) provide the incidence angles for the elements. The angle between an element and its corresponding hub axis vector defines the incidence angle. Consequently, there are two incidence angles for each element. Fig. 19a shows the incidence angles for two element centerlines connected to one hub. There is significant discontinuity between subparts A and C in this example. This discontinuity is present whenever all the elements arrive at a node from one side of the curvature of the surface. Thus, the center of extrusion of subpart A should be modified to lessen this discontinuity. Fig. 19b shows the difference between previous interconnecting part geometry and the resulting one, where the center point was modified. Fig. 19c shows the resulting geometry. To modify the center point for the extrusion of subpart A, we use the external cylinder and the lines of the mesh intersecting points, already described in Section 3.3.8.3. Fig. 20a shows the cylinder, the element centerlines, and the intersection points. These intersection points are projected on the hub axis. Then, the average of the collection of points projected on the hub axis is the new hub center point (indicated with the central arrow in Fig. 20a). Fig. 20b shows the modification of the geometries, with old and new positions, as well as the mesh center point, substituted by the average point for the generation of subpart A"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002703_j.matdes.2021.109530-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002703_j.matdes.2021.109530-Figure2-1.png",
        "caption": "Fig. 2. a) Specimen arrangement on the build plate (top view) b) 3D model of",
        "texts": [
            " The recoating mechanism consists of a rubber x-profile anddistributes a thin layer ofmetal powder from the feed region over thebuild platform. Afterwards, the applied layer is selectivelymolten according to the cross-sections of the parts andmetallurgically bonded to the underlying layer. As soon as this step is completed, the build platform is lowered and a new layer of powder is applied. These steps are repeated until the parts are finished. A standardized build process with three test specimens (cf. Fig. 2) was designed and manufactured. In order to investigate the influence of the component height on the temperature development, the first specimen represents a cylinder with a diameter of 8 mm and a height of 16 mm. In order to be able to observe up- and downskin effects, two cones were designed, which differ in the build-up process by a 180\u00b0 rotation. The cones alsomeasure 16 mm in height and have a conically reduced diameter from 8mm to 2 mm. By choosing these dimensions, the cones can bemanufacturedwithout support structures, as the critical angle of 45\u00b0 is not reached"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002821_j.engfailanal.2021.105394-Figure14-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002821_j.engfailanal.2021.105394-Figure14-1.png",
        "caption": "Fig. 14. Rotor test bench.",
        "texts": [
            " The values of index J are relatively large in the initial stage of rub-impact as shown in Fig. 13. Combining with Fig. 12, there is better stability in the index J at the initial stage of rub-impact. Meanwhile, when clearances diminish, the variations of index J become smaller and smaller. This also shows that there is a better stability in the index J used to diagnose rotor rub-impact faults and confirms that index J has better diagnosis effect for rotor system rub-impact fault than the existing fault diagnosis indexes. The experimental equipment is shown in Fig. 14. When identifying the values of NOFRFs, twice different excitations need to be loaded to the system. These excitations were obtained by changing the number of unbalanced bolts on the disk, as shown in Fig. 14. As we know, unbalanced force Fe = m\u03c92r, therefore, the force will be changed when the mass m change. During the experiment, the mass m will be changed by increasing the number of unbalanced bolts so that the vibration amplitude of the system will be changed too. This section conducted the experiments for justifying the effectiveness of index J and the correctness of the simulations. The J.T. Li et al. Engineering Failure Analysis 125 (2021) 105394 J.T. Li et al. Engineering Failure Analysis 125 (2021) 105394 experiments were performed by utilizing the test bench shown in Fig. 14, and the vibration signals were collected. Two materials (copper and aluminum alloy) were used to simulate different rub-impact stiffness and different coefficient of friction. The advantages of index J are verified by comparing the indexes changes under the twice different experiments. These devices are shown in Fig. 14, where the type of eddy current sensor is CWY \u2013 DO \u2013 502, collection card is NI \u2013 9234, and chassis of the card is Cdaq \u2013 9178, and Table 6 shows the specific parameters of the experiment. J.T. Li et al. Engineering Failure Analysis 125 (2021) 105394 The nonlinear characteristics of the rotor system were extracted by adopting the theory of WNOFRFs, and the curve Rn(2) was shown in Fig. 15. Comparing with the nonlinear characteristics caused by different materials, we can conclude that the Rn(2) change trend on the two case are roughly similar"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002752_s00170-021-06757-5-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002752_s00170-021-06757-5-Figure16-1.png",
        "caption": "Fig. 16 Grinding trajectory for the curved rake face",
        "texts": [
            " According to the theory of differential geometry, the mean curvature can be calculated by: Hi LiGi\u22122MiFi \u00fe NiEi 2 EiGi\u2212F2 i \u00f031\u00de Using Gaussian curvature and mean curvature, the maximum normal curvature is obtained by: \u03bamax \u00bc Hi \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffi Hi 2\u2212Ki p \u00f032\u00de Then, the minimum curvature radius is calculated by: \u03c1min \u00bc 1 \u03bamax \u00f033\u00de The radius of the grinding wheel should be less than the minimum curvature radius, that is Rs\u2264min \u03c1minf g \u00f034\u00de The NC program for the curved rake face is generated by CAM software after path planning and post-processing. For path planning, the parameters of the grinding wheel are set into the software according to the measured data. The steplength is calculated by equal-chord-deviation method, and the chord deviation is set as 0.001 mm. The vector direction of the wheel axis is controlled by the method of interpolation [20]. Based on the above settings, the grinding path is generated by the method of isoparametric-line [20] as shown in Fig. 16. This experiment uses Walter five-axis tool grinding machine to grind the cutter faces. The post-processing is done according to the structure and parameters of this machine to generate applicable NC program. Then, the cutter faces are ground as shown in Fig. 17 and Fig. 18. After grinding, the cutter with curved rake face (cutter A) is obtained as shown in Fig. 19. The cutter with plane rake face (cutter B) is also manufactured for comparison as shown in Fig. 20. Then, the manufacturing accuracy is measured by the measuring module of the tool grinding machine as shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002467_j.neucom.2020.09.003-Figure18-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002467_j.neucom.2020.09.003-Figure18-1.png",
        "caption": "Fig. 18. 18 The network topology G A3 .",
        "texts": [],
        "surrounding_texts": [
            "K \u00bc 0:5; 0:2; 0:3\u00f0 \u00deT , then Assumptions 2, 4 and 5 are satisfied. We chose K \u00bc I4. Based on Lemmas 4 and 5, it has kmax W\u00f0 \u00de \u00bc 4:81. For 1 \u00bc 0:3; 1 \u00bc 0:01 and 1 \u00bc 0:001, by using MATLAB software, we obtain positive definite matrices\n0B@ 1CA;\nP 0:01\u00f0 \u00de \u00bc 3:4932 1:4677 0:4913 1:4677 0:6259 0:2084 0:4913 0:2084 0:0787\n0B@ 1CA;\nP 0:001\u00f0 \u00de \u00bc 3:4721 1:4620 0:4875 1:4620 0:6165 0:2055 0:4875 0:2055 0:0694\n0B@ 1CA;\neach of which satisfies condition (12). The state trajectories of all agents and control input with different 1 are shown in the Figs. 10\u201315, respectively. We can find that all agents achieve a",
            "common state by using the low-gain control protocol (2). The amplitude of the control input does not exceed the upper bound of saturation -.\nExample 3. Consider a MAS (1) consisting of five agents where the system parameter and matrices are the same as the ones in Example 1. The network topologies are switched among Figs. 16\u201319, in which the weight of each edge is 1. The initial states of all agents\nare randomly selected from a bounded set 1;1\u00bd 1;1\u00bd 1;1\u00bd . Choosing K \u00bc 0:5; 0:3; 0:4\u00f0 \u00deT , then Assumptions 2, 4 and 6 are satisfied. In the simulation, we choose K. ~t1\u00f0 \u00de \u00bc K. ~t2\u00f0 \u00de \u00bc 0:2I4, then c \u00bc 0:2 and c \u00bc 0:96.\nFor 1 \u00bc 0:1, by using MATLAB software, we obtain a positive definite matrix\nP 0:1\u00f0 \u00de \u00bc 0:5036 0:4203 0:0699 0:4203 0:4349 0:0658 0:0699 0:0658 0:0706\n0B@ 1CA",
            "which satisfies condition (23). By calculating, we get that g1 \u00bc 0:11;g3 \u00bc 56:2. Then, we chose the communication intervals to satisfy condition (24). The state trajectories of all agents under the protocol (2) are shown in Fig. 20. The control input is presented in Fig. 21.\nFor 1 \u00bc 0:01, by using MATLAB software, we obtain a positive definite matrix\nP 0:01\u00f0 \u00de \u00bc 0:1171 0:1034 0:0156 0:1034 0:1000 0:0145 0:0156 0:0145 0:0082\n0B@ 1CA\nwhich satisfies condition (23). By calculating, we get that g1 \u00bc 0:046;g3 \u00bc 19:1. Then, we chose the communication intervals to satisfy condition (24). The state trajectories of all agents under the protocol (2) are shown in Fig. 22. The control input is presented in Fig. 23.\nIt can be seen from Figs. 20\u201322 that all agents reach consensus under protocol (2) with different 1 and the amplitude of the control input does not exceed the upper bound of saturation -.\n7. Conclusion\nIn this paper, we investigate the consensus problem of MASs with input saturation on directed networks. Firstly, we study the leaderless consensus of MASs over directed strongly connected networks. Secondly, based on tree-type error method, the consensus of MASs with directed spanning tree is investigated. The stability is proved by using the Hurwitz stability and Lyapunov stability theorem. In addition, the input saturation consensus on general switching networks is considered, in which the ratio of the communication width between the topology with and without a directed spanning tree is obtained. In the network communication, the network-induced effects are very important for the consensus of MASs. Therefore, the distributed consensus subject to quantization effects [49] and protocol scheduling effects [50] will be considered in our future work.\nCRediT authorship contribution statement\nShuzhen Yu: Writing - review & editing. Zhiyong Yu: Data curation, Writing - original draft. Haijun Jiang: Supervision, Conceptualization, Methodology. Xuehui Mei: Software, Validation.\nDeclaration of Competing Interest\nThe authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.\nReferences\n[1] J. Fax, R. Murray, Information flow and cooperative control of vehicle formations, IEEE Trans. Autom. Control 49 (1) (2004) 115\u2013120. [2] W. Yu, G. Chen, Z. Wang, W. Yang, Distributed consensus filtering in sensor networks, IEEE Trans. Syst. Man Cybern. B Cybern. 39 (6) (2009) 1568\u20131577. [3] W. Ren, R. Beard, Consensus seeking in multiagent systems under dynamically changing interaction topologies, IEEE Trans. Autom. Control 50 (5) (2005) 655\u2013 661. [4] J. Berni, P. Zarcotejada, L. Suarez, E. Fereres, Thermal and narrowband multispectral remote sensing for vegetation monitoring from an unmanned aerial vehicle, IEEE Trans. Geosci. Remote Sens. 47 (3) (2009) 722\u2013738. [5] J. Mei, W. Ren, G. Ma, Distributed coordinated tracking with a dynamic leader for multiple euler-lagrange systems, IEEE Trans. Autom. Control 56 (6) (2011) 1415\u20131421. [6] L. Ying, G. Yan, Z. Lin, Synthesis of distributed control of coordinated path following based on hybrid approach, IEEE Trans. Autom. Control 56 (5) (2011) 1170\u20131175. [7] H. Min, F. Sun, S. Wang, H. Li, Distributed adaptive consensus algorithm for networkedEuler-Lagrangesystems, IETControlTheoryAppl.5 (1) (2011)145\u2013154. [8] X. Liu, A. Aichhorn, L. Liu, H. Li, Coordinated control of distributed energy storage systemwith tap changer transformers for voltage rise mitigation under high photovoltaic penetration, IEEE Trans. Smart Grid 3 (2) (2012) 897\u2013906. [9] K. Tan, P. So, Y. Chu, M. Chen, Coordinated control and energy management of distributed generation inverters in a microgrid, IEEE Trans. Power Delivery 28 (2) (2013) 704\u2013713. [10] R. Olfati-Saber, R. Murray, Consensus problems in networks of agents with switching topology and time-delays, IEEE Trans. Autom. Control 49 (9) (2004) 1520\u20131533. [11] W. Ren, E. Atkins, Distributed multi-vehicle coordinated control via local information exchange, Int. J. Robust Nonlinear Control 17 (10\u201311) (2010) 1002\u20131033. [12] G. Xie, L. Wang, Consensus control for a class of networks of dynamic agents, Int. J. Robust Nonlinear Control 17 (10\u201311) (2007) 941\u2013959. [13] J. Wang, D. Cheng, X. Hu, Consensus of multi-agent linear dynamic systems, Asian J. Control 10 (2) (2010) 144\u2013155. [14] Z. Yu, H. Jiang, D. Huang, C. Hu, Consensus of nonlinear multi-agent systems with directed switching graphs: a directed spanning tree based error system approach, Nonlinear Anal. Hybrid Syst. 28 (2018) 123\u2013140. [15] Z. Yu, H. Jiang, C. Hu, X. Fan, Consensus of second-order multi-agent systems with delayed nonlinear dynamics and aperiodically intermittent communications, Int. J. Control 90 (5) (2016) 909\u2013922. [16] Z. Yu, H. Jiang, D. Huang, C. Hu, Directed spanning tree based adaptive protocols for second-order consensus of multiagent systems, Int. J. Robust Nonlinear Control 28 (6) (2018) 2172\u20132190. [17] Y. Qian, X. Wu, J. L\u00fc, J. Lu, Second-order consensus of multi-agent systems with nonlinear dynamics via impulsive control, Neurocomputing 125 (2014) 142\u2013 147. [18] S. Yoo, Distributed consensus tracking for multiple uncertain nonlinear strictfeedback systems under a directed graph, IEEE Trans. Neural Networks Learn. Syst. 24 (4) (2013) 666\u2013672. [19] Z. Li, X. Liu, P. Lin, W. Ren, Consensus of linear multi-agent systems with reduced-order observer-based protocols, Syst. Control Lett. 60 (7) (2011) 510\u2013 516. [20] C. Yoshioka, T. Namerikawa, Observer-based consensus control strategy for multi-agent system with communication time delay, in: 2008 IEEE International Conference on Control Applications, San Antonio, TX, 2008, pp. 1037\u20131042. [21] D. Ding, Z. Wang, W. Daniel, G. Wei, Observer-based event-triggering consensus control for multiagent systems with lossy sensors and cyberattacks, IEEE Trans. Cybern. 47 (8) (2017) 1936\u20131947. [22] Y. Cao, L. Zhang, C. Li, Z. Michael, Observer-based consensus tracking of nonlinear agents in hybrid varying directed topology, IEEE Trans. Cybern. 47 (8) (2017) 2212\u20132222. [23] Z. Yu, S. Yu, H. Jiang, C. Hu, Observer-based consensus for multi-agent systems with partial adaptive dynamic protocols, Nonlinear Anal. Hybrid Syst. 34 (2019) 58\u201373. [24] L. Rong, S. Wang, G. Jiang, S. Xu, Distributed observer-based consensus over directed networks with limited communication bandwidth constraints, IEEE Trans. Syst. Man Cybern. Syst. doi. DOI: 10.1109/TSMC.2018.2875523. [25] L. Jian, J. Hu, J. Wang, K. Shi, Observer-based output feedback distributed event-triggered control for linear multi-agent systems under general directed graphs, Physica A 534 (2019) 122288. [26] M. Ran, Y. Hou, X. Gong, Q. Wang, C. Dong, Distributed output-feedback consensus control of multi-agent systems with dynamically changing directed interaction topologies, ISA Trans. 85 (2019) 71\u201375. [27] J. Xu, Y. Tan, T. Lee, Iterative learning control design based on composite energy function with input saturation, Automatica 40 (8) (2004) 1371\u20131377. [28] Y. Tang, Z. Wang, H. Gao, H. Qiao, J. Kurths, On controllability of neuronal networks with constraints on the average of control gains, IEEE Trans. Cybern. 44 (12) (2014) 2670\u20132681. [29] H. Su, G. Jia, M. Chen, Semi-global containment control of multi-agent systems with intermittent input saturation, J. Franklin Inst. 352 (9) (2015) 3504\u20133525. [30] Z. Lin, Global control of linear systems with saturating actuators, Automatica 34 (7) (1998) 897\u2013905."
        ]
    },
    {
        "image_filename": "designv11_14_0002501_j.engfailanal.2020.104977-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002501_j.engfailanal.2020.104977-Figure1-1.png",
        "caption": "Fig. 1. Failures of elastic elements of flexible coupling.",
        "texts": [
            " That is why the maximum operating temperature is limited at 70 to 110 \u25e6C, depending on material. [11\u201313]. Heating of flexible members leads to changes in dynamic properties of couplings [14,15] and drivetrain failures [16,17]. It is therefore necessary to investigate the effect of external dynamic loading on heat generation in flexible couplings and thereby eliminate one of the potential causes of failure. We prove our claim with practical demonstrations of elastic coupling elements failures. In Fig. 1, there are three different types of flexible couplings (A) B-flex RB 116-4, (B) Periflex PNA 10R, (C) Gurimax GVW 100. Torsional vibration of elastic couplings caused self-heating of flexible elements and consequently the increased temperature caused the failure of elastic elements. The experiments were carried out at ambient temperature of 20 \u25e6C. Failures of the elastic elements are indicated by arrows [18]. The use of air and its properties is a promising option. Pneumatic actuation systems are used throughout the industry whenever economic, low-maintenance, medium-range force, and simple-motion solutions must be implemented"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003913_ip-cta:19990588-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003913_ip-cta:19990588-Figure1-1.png",
        "caption": "Fig. 1",
        "texts": [
            " The use of the reduced-order observer also simplifies the comparison between the design in [12] and the FI H , approach. An alternative approach based on H , output feedback is more complex and would yield a high order controller that can also become unstable as we approach the optimal design. From a practical point of view an unstable controller can cause substantial actuator saturation and could fail to stabilise the triple pendulum system. This alternative is left for future research. 2 System formulation The schematic representation of the triple inverted pendulum system considered here shown in Fig. 1. Ll , L2, L3 are the lengths of the lower, middle and upper link respectively; 01, Q2, O3 are the angles of the lower, middle, and upper link with respect to the vertical line; Y is the cart position; a is the rail inclination; 71, T ~ , z3 are disturbance torques at the lower, middle and upper link respectively; a i , a2, a3 denote the position of center of mass of lower link, middle link, and third link; F, is the dry friction along the rail, and U the input voltage to the motor. In addition to the input voltage of the motor, other external disturbances affect the system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000784_s12206-016-1123-4-Figure10-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000784_s12206-016-1123-4-Figure10-1.png",
        "caption": "Fig. 10. Schematic of gear test rig.",
        "texts": [
            " This result shows that it is necessary to careful apply a condition indicator, such as a crest factor or kurtosis, to analyze the impulse character of a vibration signal, when defining the level of the fault condition in a vibration-based condition monitoring system. It also indicates that it is difficult to apply a constant alarm or fault level in a vibration-based monitoring system for a gearbox in railway vehicles, where there is a successive change of torque during operation. To analyze the effects of the damaged gear upon the vibration of gearbox, gear dynamic tests are carried out using the gear test rig illustrated in Fig. 10. The test system consists of an input motor that controls rotational velocity, an output motor that controls torque, and a motor control interface and gear specimen situated between the input and output motor. Accelerometers to measure gear vibrations are attached onto the bearing housing nearest to the gear specimen. A sampling rate of 10.2 kS/s was chosen for vibration data acquisition, and Fig. 11 shows the vibration signal in a time domain for 2 seconds for undamaged and damaged gear (full-tooth fracture of teeth) at 500 rpm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000212_mees.2019.8896501-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000212_mees.2019.8896501-Figure5-1.png",
        "caption": "Fig. 5. Radial AMB models: a) the geometric one in the air; b) the geometric one of the stator, rotor and coils; c) the finite-element one.",
        "texts": [
            " Experimental data were obtained using a laboratory rig [5, 6] by measuring the forces, that balance the rotor in a certain position, with a dynamometer. The solid unmarked lines in Fig. 3 are an approximation of the calculated data by a cubic polynomial. Analysis of the results indicated that a \u201ccalculation-experiment\u201d discrepancy does not exceed 1%. This proves an applicability and accuracy of the technique. The block diagram of one of the CS variants for a radial AMB is shown in Fig. 4 [9]. Geometric and finite element models of the radial AMB with an eight-pole stator are presented in Fig. 5. Here the fixed Cartesian coordinate system is entered. The control system (Fig. 4) implements the control law. This law determines the currents in four electromagnets (ic1,...,ic4) depending on a spatial position of the rotor, taking into account displacement currents that create a force tension. The difference in the method for calculating the dependences of magnetic forces on a rotor displacement for a radial AMB is as follows. The full nominal gap 2 r (on both sides of a collar) is evenly divided into (2n+1) levels in both the vertical and horizontal directions"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000276_s00366-019-00893-z-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000276_s00366-019-00893-z-Figure1-1.png",
        "caption": "Fig. 1 A sliding cable with N \u2212 1 sub-elements",
        "texts": [
            "\u00a02 introduces the basic equations governing the behavior of sliding cables. In Sect.\u00a03, sliding calculations through LC approach are presented. Section\u00a04 introduces the LCenhanced transient stiffness method and dynamic relaxation method. The proposed formulations are then explored and compared through numerical examples in Sect.\u00a04. Finally, conclusions are discussed in Sect.\u00a05. Consider a continuous cable that connects N nodes. The cable can be decomposed of N \u2212 1 sub-elements, while sliding over N \u2212 2 pulleys that are assumed attached to its intermediate nodes (Fig.\u00a01). The cable is assumed to be tensioned. The length of the sliding cable is given by the sum of all sub-element lengths: where lj is the length of the sub-element ej. Equivalently, the rest-length (free/unstretched length) of the sliding element, marked with the subscript zero, is defined as: where l0,j is the rest-length of the sub-elements ej. If the continuous cable is assumed to run over frictionless pulleys, then the tensile force t acting on all sub-elements is given by: (1)l = N\u22121\u2211 j=1 lj, (2)l0 = N\u22121\u2211 j=1 l0,j 1 3 where E and A are the Young\u2019s modulus and the cross-section area of the sliding cable",
            " The above formulation reflects the basic workhorse for most existing finite element formulations [4, 8, 10, 12, 33] and dynamic relaxation implementations [18]. As already explained, ignoring friction on sliding nodes may result in substantial model errors as with friction the tension forces at different sub-elements are no longer the same and, thus, cannot be obtained based on nodal positions only. For a more accurate modeling of the static behavior of tensile structures with sliding cables, friction at sliding nodes must thus be considered. To understand the effect of sliding-induced friction, consider the structure shown in Fig.\u00a01. Assume that friction exists at the contact area between the pulley and the cable at node i. In such situation the relative motion between cable sub-elements ei\u22121 and ei is affected by friction. Cable sliding is possible when friction forces are low. However, reaching a certain level, friction forces impede the cable from sliding. If this occurs, the internal force ti in the sub-element ei can be expressed as: where li is the deformed length of the sub-element ei, while l\u03040,i is its corresponding rest-length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002572_j.apm.2020.12.020-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002572_j.apm.2020.12.020-Figure5-1.png",
        "caption": "Fig. 5. Schematic diagram of tracking mechanism.",
        "texts": [
            " where F \u2032 T e , p = [ 0 0 \u2212m E g 0 0 0 ] . It can be seen that the dynamic model of the tracking mechanism manifested in Eq. (35) is achieved through assembling the dynamic model of an arbitrary flexible body j obtained via virtually cutting joints of the mechanism, which contains no constraints. Consequently, in order to establish the complete dynamic model of the system, some constraint equations of the system should be introduced into the open-loop dynamic model. The schematic diagram of the tracking mechanism is shown in Fig. 5 . The mechanism is divided into 17 flexible bodies, each of which uses a spatial beam element, totaling 17 elements. As the joint position coordinates of flexible bodies are introduced in the derivation of dynamic equations of arbitrary flexible bodies, the x -axis of joint coordinates of flexible bodies is along the length direction of each limb, and the y -axis is parallel to the axis of the rotational joint connected with the limb before deformation, and the z -axis satisfies the right-hand rule"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000298_s12555-018-0946-4-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000298_s12555-018-0946-4-Figure1-1.png",
        "caption": "Fig. 1. The underactuated AUV model in inertial and body fixed frames.",
        "texts": [
            " Statics is the stability of the device in steady or moving at constant velocity states and the dynamics is related to the accelerated motion. Usually, dynamic studies are divided into two parts: kinematics that only examines the geometric dimension of motion and kinetics that analyzes the forces that generate motion [23]. Modeling and describing the function of the system by a series of dynamic equations based on which the controller is designed is very important. Given the activities carried out in the field of AUV modeling, a relatively comprehensive model of underwater vehicle has been extracted [24, 25]. Fig. 1 presents an example of AUV and its dynamic variables in the body-fixed coordinate frame and its position relative to the inertial coordinate frame. Table 1 indicates different model variables defined as AUV\u2019s motion behaviors in accordance with the Society of Naval Architects and Marine Engineers (SNAME). 2.1. AUV\u2019s dynamic equations In this section, nonlinear dynamic equations are extracted to design the controller in the depth and steer motion. To create a dynamic model for AUV, first, it is necessary to obtain kinetic equations and then analyze and extract external forces in the equation of motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001531_1.c035829-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001531_1.c035829-Figure4-1.png",
        "caption": "Fig. 4 UAV force analysis diagram.",
        "texts": [
            " The model of the tire ground reaction forces Pn, Pml, and Pmr can be regarded as a first-order spring-damper system [30]. In the ground axis system, the reaction forces are given as follows: Pnm 0 0 \u2212 Pn Pml Pmr T (21) And in Sb \u2212Obxbybzb, the corresponding moments are Mp 2 64 Pml \u2212 Pmr \u22c5 bw\u22152 \u22c5 cos\u03d5 Pnan cos \u03b8 \u2212 Pml Pmr am cos \u03b8 0 3 75 (22) where \u03d5 and \u03b8 are the roll and pitch angles of the UAV. A 6-DOF UAV ground taxiing mathematical model is built in this study, and the UAV force analysis diagram is illustrated in Fig. 4. In addition to the forces calculated in Secs. II.A\u2013II.D, the gravity G, the motor power Ft, and the ground friction fx on the tires still need to be taken into consideration. The tire longitudinal friction can be given in Ss \u2212Osxsyszs: fx \u2212 fxml fxmr fxn cos \u03b8l fxn sin \u03b8l 0 T (23) where fxn, fxml, and fxmr are the tire longitudinal friction forces on the nose and twomain wheels. And the moments of the friction to the UAVare expressed in Sb \u2212Obxbybzb: Mfx 2 664 fxn sin \u03b8lhn \u2212fxn cos \u03b8lhn \u2212 fxmlhml fxmrhmr \u2212fxn sin \u03b8lan \u2212fxml fxmr \u22c5 bw\u22152 3 775 (24) According to the momentum theorem, the mass-center dynamic vector equations of the UAV in a moving coordinate system can be derived as M \u03b4V \u03b4t \u03c9 \u00d7 V F (25) where \u03c9 is the rotational speed of the moving coordinate system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002268_012047-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002268_012047-Figure3-1.png",
        "caption": "Figure 3. The cross section shear band of (a) Tweel airless tire and (b) airless tire model.",
        "texts": [
            " The experimental result was indicated that the vertical stiffness of airless tire which was studied is 949.37 N/mm. The International Conference on Materials Research and Innovation (ICMARI) IOP Conf. Series: Materials Science and Engineering 773 (2020) 012047 IOP Publishing doi:10.1088/1757-899X/773/1/012047 The computer aided design (CAD) and computer aided engineering (CAE) was use to create the 3D finite element model of Tweel airless tire. The finite element model of steel belt layer is developed according to the Tweel airless tire which has a complex shear band construction as shown in figure 3. The shear band of finite element model composed wall and belt elements. The tread material property of airless tire model was investigated by compresion test according to ASTM D575 standard, while material property of spoke was investigated by tensile test according to ASTM D412 standard, respectively. The tread, wall and spoke properties were defined by Mooney-Rivlin hyperelastic model. On the other hand, the steel belt is specified to be an elastic isotropic material as described in table 2. The airless tire model was compressed with the compressive load of 1,000 kg"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000464_j.isatra.2015.12.017-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000464_j.isatra.2015.12.017-Figure4-1.png",
        "caption": "Fig. 4. Four-DOF bearing-pedestal model.",
        "texts": [
            " If the rolling element is not located in the defected region, the actual radial clearance ra is given by ra \u00bc r0: \u00f013\u00de If the rolling element is located in the defected region, the actual radial clearance ra is given by ra \u00bc r0\u00fe\u0394d: \u00f014\u00de Therefore, the contact forces can be calculated as follows: Fbx \u00bc XN i \u00bc 1 \u03b3iKb x cos\u03b8i\u00fey sin\u03b8i ra 1:5 cos\u03b8i; \u00f015\u00de Fby \u00bc XN i \u00bc 1 \u03b3iKb x cos \u03b8i\u00fey sin\u03b8i ra 1:5 sin \u03b8i: \u00f016\u00de Since the rotating speed is far below the first bending critical speed, the rotor is modeled as a rigid rotor and the gyroscopic effect is neglected. The model of Feng et al. [8] is adopted to study the dynamics of rolling element bearings, which is shown in Fig. 4. The model has four DOF, including two DOF of inner race (xs,ys) and two DOF of pedestal (xp, yp). s exciters for rolling element bearing outer race defect diagnosis. i The total contact forces of the model considering inner race and pedestal can be calculated as follows: Fbx \u00bc XN i \u00bc 1 \u03b3iKb xs xp cos\u03b8i\u00fe ys yp sin\u03b8i ra h i1:5 cos\u03b8i; \u00f017\u00de Fby \u00bc XN i \u00bc 1 \u03b3iKb xs xp cos\u03b8i\u00fe ys yp sin\u03b8i ra h i1:5 sin\u03b8i: \u00f018\u00de Considering the nonlinear contact force, the governing equations of motion can be written as follows: ms \u20acxs\u00fecs _xs\u00feFbx \u00bc Fu cos\u03c9st\u00feFMX ; \u00f019\u00de ms \u20acys\u00fecs _ys\u00feFby \u00bc Fu sin\u03c9st msg\u00feFMY ; \u00f020\u00de mp \u20acxp\u00fecp _xp\u00feKpxp Fbx \u00bc 0; \u00f021\u00de mp \u20acyp\u00fecp _yp\u00feKpyp Fby \u00bc mpg; \u00f022\u00de where ms is the mass of the shaft and the inner race; cs is the bearing damping; mp is the mass of the pedestal; Kp and cp are the stiffness and damping of the pedestal"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003715_s0301-679x(98)00106-6-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003715_s0301-679x(98)00106-6-Figure8-1.png",
        "caption": "Fig. 8. Scheme of the transmission photoelastic rig.",
        "texts": [
            " The seal segment and the thin photoelastic sheet are schematically represented; the photoelastic material is fastened to the seal section. The photoelastic sheet is made of special material for measurements on rubber; it is the PS-6 Sheet and the glue is the PC-9 Adhesive of Measurement Group. The Young\u2019s modulus of this material is E=0.7 MPa; this value is lower than the Young\u2019s modulus of the rubber under test (E of about 10 MPa). In order to have a better visualisation of isoclinic lines a transparent photoelastic rig has been developed. The rig, schematically shown in Fig. 8, has two glass plates fixed to the frame. The gap between the two plates can be regulated by precision screws. Tests have been carried out using a scale model of the seal with 3.5\u00d7 magnification. The material employed to make the scale model is a low stiffness photoelastic sheet. The seal model is mounted in a housing made of steel; under the seal there is a plate simulating the cylinder rod. The working conditions are related to the seal mounting without fluid pressure and without rod reciprocating motion"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001193_j.ijsolstr.2019.09.010-Figure23-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001193_j.ijsolstr.2019.09.010-Figure23-1.png",
        "caption": "Figure 23: (a) Initial setup of tape spring coiling simulations in Section 4.1; (b) associated dimensions (in mm) and node sets.",
        "texts": [
            " The displacement boundary conditions ux,fold and uy,fold were imposed to coil the tape spring through an angle \u03b8: ux,fold =rs(1\u2212 cos \u03b8) + (LB \u2212 rs\u03b8) sin \u03b8 uy,fold =rs sin \u03b8 + (LB \u2212 rs\u03b8) cos \u03b8 \u2212 LB (19) Contact with the temporary cylinders was then removed and the tape spring was allowed to find an equilibrium configuration. Mass nodal damping was335 applied to remove any excess kinetic energy. This simulation technique resulted in a two-fold configuration, as shown in Fig. 14(b-c). Details of the applied boundary conditions, contact conditions, and corresponding node sets are given in Table 3 and Fig. 23 in the Appendix. As the extension ux of the tape spring was increased, it was critical to340 determine at which stage of the simulation the tape spring would be considered to be fully coiled. The fully coiled configuration was defined such that the centerline of the tape spring conformed to the cylinder surface to within three tape spring thicknesses over a region defined by \u03b3min \u2264 \u03b3 \u2264 \u03b3max. The range \u03b3min, \u03b3max was chosen to remove localized effects near the the clamped end and345 the first point of contact between the tape spring and the cylinder"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003527_pime_proc_1938_139_010_02-Figure30-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003527_pime_proc_1938_139_010_02-Figure30-1.png",
        "caption": "Fig. 30. Changes in a Fluid Coupling when Accelerator is Released",
        "texts": [
            " Thus the action of releasing the accelerator causes a reversal in the torque transmitted through the fluid coupling, and effects the change of ratio in the Cotal gearbox at the same instant, i.e. when the liquid vortex ring is in a collapsed condition and momentarily incapable of transmitting high torque. In other words, the vortex behaves at this instant as if the coupling were largely emptied of liquid. The sequence of changes in the fluid coupling when the accelerator is released, so that the torque changes from the driving to the overrunning condition, is illustrated by Fig. 30. The normal driving condition with clockwise cir- 2016 at UNIV CALIFORNIA SANTA BARBARA on March 13,pme.sagepub.comDownloaded from PROBLEMS OF FLUID COUPLINGS 125 culation of the liquid in the vortex is shown by Fig. 30 a. When the engine power is cut off the vortex ring collapses as shown at b and for perhaps half a second it consists of an eddying confusion of liquid, mixed with any air present. Half a second later the vortex circulation re-forms in the anti-clockwise direction as shown at c and the fluid coupling is again capable of transmitting high torque. The control switch with its interlock is so arranged that it is possible to change down by preselecting the required gear and pushing the accelerator pedal down to the limit of its travel, so as to trip the switch blade and make the gear change with the engine at full throttle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002467_j.neucom.2020.09.003-Figure19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002467_j.neucom.2020.09.003-Figure19-1.png",
        "caption": "Fig. 19. The network topology G A4 .",
        "texts": [],
        "surrounding_texts": [
            "K \u00bc 0:5; 0:2; 0:3\u00f0 \u00deT , then Assumptions 2, 4 and 5 are satisfied. We chose K \u00bc I4. Based on Lemmas 4 and 5, it has kmax W\u00f0 \u00de \u00bc 4:81. For 1 \u00bc 0:3; 1 \u00bc 0:01 and 1 \u00bc 0:001, by using MATLAB software, we obtain positive definite matrices\n0B@ 1CA;\nP 0:01\u00f0 \u00de \u00bc 3:4932 1:4677 0:4913 1:4677 0:6259 0:2084 0:4913 0:2084 0:0787\n0B@ 1CA;\nP 0:001\u00f0 \u00de \u00bc 3:4721 1:4620 0:4875 1:4620 0:6165 0:2055 0:4875 0:2055 0:0694\n0B@ 1CA;\neach of which satisfies condition (12). The state trajectories of all agents and control input with different 1 are shown in the Figs. 10\u201315, respectively. We can find that all agents achieve a",
            "common state by using the low-gain control protocol (2). The amplitude of the control input does not exceed the upper bound of saturation -.\nExample 3. Consider a MAS (1) consisting of five agents where the system parameter and matrices are the same as the ones in Example 1. The network topologies are switched among Figs. 16\u201319, in which the weight of each edge is 1. The initial states of all agents\nare randomly selected from a bounded set 1;1\u00bd 1;1\u00bd 1;1\u00bd . Choosing K \u00bc 0:5; 0:3; 0:4\u00f0 \u00deT , then Assumptions 2, 4 and 6 are satisfied. In the simulation, we choose K. ~t1\u00f0 \u00de \u00bc K. ~t2\u00f0 \u00de \u00bc 0:2I4, then c \u00bc 0:2 and c \u00bc 0:96.\nFor 1 \u00bc 0:1, by using MATLAB software, we obtain a positive definite matrix\nP 0:1\u00f0 \u00de \u00bc 0:5036 0:4203 0:0699 0:4203 0:4349 0:0658 0:0699 0:0658 0:0706\n0B@ 1CA",
            "which satisfies condition (23). By calculating, we get that g1 \u00bc 0:11;g3 \u00bc 56:2. Then, we chose the communication intervals to satisfy condition (24). The state trajectories of all agents under the protocol (2) are shown in Fig. 20. The control input is presented in Fig. 21.\nFor 1 \u00bc 0:01, by using MATLAB software, we obtain a positive definite matrix\nP 0:01\u00f0 \u00de \u00bc 0:1171 0:1034 0:0156 0:1034 0:1000 0:0145 0:0156 0:0145 0:0082\n0B@ 1CA\nwhich satisfies condition (23). By calculating, we get that g1 \u00bc 0:046;g3 \u00bc 19:1. Then, we chose the communication intervals to satisfy condition (24). The state trajectories of all agents under the protocol (2) are shown in Fig. 22. The control input is presented in Fig. 23.\nIt can be seen from Figs. 20\u201322 that all agents reach consensus under protocol (2) with different 1 and the amplitude of the control input does not exceed the upper bound of saturation -.\n7. Conclusion\nIn this paper, we investigate the consensus problem of MASs with input saturation on directed networks. Firstly, we study the leaderless consensus of MASs over directed strongly connected networks. Secondly, based on tree-type error method, the consensus of MASs with directed spanning tree is investigated. The stability is proved by using the Hurwitz stability and Lyapunov stability theorem. In addition, the input saturation consensus on general switching networks is considered, in which the ratio of the communication width between the topology with and without a directed spanning tree is obtained. In the network communication, the network-induced effects are very important for the consensus of MASs. Therefore, the distributed consensus subject to quantization effects [49] and protocol scheduling effects [50] will be considered in our future work.\nCRediT authorship contribution statement\nShuzhen Yu: Writing - review & editing. Zhiyong Yu: Data curation, Writing - original draft. Haijun Jiang: Supervision, Conceptualization, Methodology. Xuehui Mei: Software, Validation.\nDeclaration of Competing Interest\nThe authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.\nReferences\n[1] J. Fax, R. Murray, Information flow and cooperative control of vehicle formations, IEEE Trans. Autom. Control 49 (1) (2004) 115\u2013120. [2] W. Yu, G. Chen, Z. Wang, W. Yang, Distributed consensus filtering in sensor networks, IEEE Trans. Syst. Man Cybern. B Cybern. 39 (6) (2009) 1568\u20131577. [3] W. Ren, R. Beard, Consensus seeking in multiagent systems under dynamically changing interaction topologies, IEEE Trans. Autom. Control 50 (5) (2005) 655\u2013 661. [4] J. Berni, P. Zarcotejada, L. Suarez, E. Fereres, Thermal and narrowband multispectral remote sensing for vegetation monitoring from an unmanned aerial vehicle, IEEE Trans. Geosci. Remote Sens. 47 (3) (2009) 722\u2013738. [5] J. Mei, W. Ren, G. Ma, Distributed coordinated tracking with a dynamic leader for multiple euler-lagrange systems, IEEE Trans. Autom. Control 56 (6) (2011) 1415\u20131421. [6] L. Ying, G. Yan, Z. Lin, Synthesis of distributed control of coordinated path following based on hybrid approach, IEEE Trans. Autom. Control 56 (5) (2011) 1170\u20131175. [7] H. Min, F. Sun, S. Wang, H. Li, Distributed adaptive consensus algorithm for networkedEuler-Lagrangesystems, IETControlTheoryAppl.5 (1) (2011)145\u2013154. [8] X. Liu, A. Aichhorn, L. Liu, H. Li, Coordinated control of distributed energy storage systemwith tap changer transformers for voltage rise mitigation under high photovoltaic penetration, IEEE Trans. Smart Grid 3 (2) (2012) 897\u2013906. [9] K. Tan, P. So, Y. Chu, M. Chen, Coordinated control and energy management of distributed generation inverters in a microgrid, IEEE Trans. Power Delivery 28 (2) (2013) 704\u2013713. [10] R. Olfati-Saber, R. Murray, Consensus problems in networks of agents with switching topology and time-delays, IEEE Trans. Autom. Control 49 (9) (2004) 1520\u20131533. [11] W. Ren, E. Atkins, Distributed multi-vehicle coordinated control via local information exchange, Int. J. Robust Nonlinear Control 17 (10\u201311) (2010) 1002\u20131033. [12] G. Xie, L. Wang, Consensus control for a class of networks of dynamic agents, Int. J. Robust Nonlinear Control 17 (10\u201311) (2007) 941\u2013959. [13] J. Wang, D. Cheng, X. Hu, Consensus of multi-agent linear dynamic systems, Asian J. Control 10 (2) (2010) 144\u2013155. [14] Z. Yu, H. Jiang, D. Huang, C. Hu, Consensus of nonlinear multi-agent systems with directed switching graphs: a directed spanning tree based error system approach, Nonlinear Anal. Hybrid Syst. 28 (2018) 123\u2013140. [15] Z. Yu, H. Jiang, C. Hu, X. Fan, Consensus of second-order multi-agent systems with delayed nonlinear dynamics and aperiodically intermittent communications, Int. J. Control 90 (5) (2016) 909\u2013922. [16] Z. Yu, H. Jiang, D. Huang, C. Hu, Directed spanning tree based adaptive protocols for second-order consensus of multiagent systems, Int. J. Robust Nonlinear Control 28 (6) (2018) 2172\u20132190. [17] Y. Qian, X. Wu, J. L\u00fc, J. Lu, Second-order consensus of multi-agent systems with nonlinear dynamics via impulsive control, Neurocomputing 125 (2014) 142\u2013 147. [18] S. Yoo, Distributed consensus tracking for multiple uncertain nonlinear strictfeedback systems under a directed graph, IEEE Trans. Neural Networks Learn. Syst. 24 (4) (2013) 666\u2013672. [19] Z. Li, X. Liu, P. Lin, W. Ren, Consensus of linear multi-agent systems with reduced-order observer-based protocols, Syst. Control Lett. 60 (7) (2011) 510\u2013 516. [20] C. Yoshioka, T. Namerikawa, Observer-based consensus control strategy for multi-agent system with communication time delay, in: 2008 IEEE International Conference on Control Applications, San Antonio, TX, 2008, pp. 1037\u20131042. [21] D. Ding, Z. Wang, W. Daniel, G. Wei, Observer-based event-triggering consensus control for multiagent systems with lossy sensors and cyberattacks, IEEE Trans. Cybern. 47 (8) (2017) 1936\u20131947. [22] Y. Cao, L. Zhang, C. Li, Z. Michael, Observer-based consensus tracking of nonlinear agents in hybrid varying directed topology, IEEE Trans. Cybern. 47 (8) (2017) 2212\u20132222. [23] Z. Yu, S. Yu, H. Jiang, C. Hu, Observer-based consensus for multi-agent systems with partial adaptive dynamic protocols, Nonlinear Anal. Hybrid Syst. 34 (2019) 58\u201373. [24] L. Rong, S. Wang, G. Jiang, S. Xu, Distributed observer-based consensus over directed networks with limited communication bandwidth constraints, IEEE Trans. Syst. Man Cybern. Syst. doi. DOI: 10.1109/TSMC.2018.2875523. [25] L. Jian, J. Hu, J. Wang, K. Shi, Observer-based output feedback distributed event-triggered control for linear multi-agent systems under general directed graphs, Physica A 534 (2019) 122288. [26] M. Ran, Y. Hou, X. Gong, Q. Wang, C. Dong, Distributed output-feedback consensus control of multi-agent systems with dynamically changing directed interaction topologies, ISA Trans. 85 (2019) 71\u201375. [27] J. Xu, Y. Tan, T. Lee, Iterative learning control design based on composite energy function with input saturation, Automatica 40 (8) (2004) 1371\u20131377. [28] Y. Tang, Z. Wang, H. Gao, H. Qiao, J. Kurths, On controllability of neuronal networks with constraints on the average of control gains, IEEE Trans. Cybern. 44 (12) (2014) 2670\u20132681. [29] H. Su, G. Jia, M. Chen, Semi-global containment control of multi-agent systems with intermittent input saturation, J. Franklin Inst. 352 (9) (2015) 3504\u20133525. [30] Z. Lin, Global control of linear systems with saturating actuators, Automatica 34 (7) (1998) 897\u2013905."
        ]
    },
    {
        "image_filename": "designv11_14_0003023_s11071-021-06591-0-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003023_s11071-021-06591-0-Figure8-1.png",
        "caption": "Fig. 8 A simplified overlooking diagram of the head-yawing",
        "texts": [
            " Based on the momentum theorem, the increased momentum of fluid caused by the fish head-yawing should be as same as the increased RF momentum from the obstructive thrust within the sampling period Dt. Accordingly, the dynamic model can be given by: TODt \u00bc Mhvh \u00f05\u00de where Mh is the fluid mass swept by the head-yawing within Dt, and vh is the longitudinal swing speed of the fish head. Considering the traditional fish body wave function, it can be seen that the swing amplitude of the fish body becomes zero as the envelope line intersects at the origin of coordinates, which indicates that there is no oscillating at this point set as O in Fig. 8 during an oscillating period of fish body. Figure 8 shows the simplified diagram of the head-yawing. In this work, the RF is designed as a pedestal-free crank-connecting system, which means equal swing angles at both sides of point O can be obtained as follows: hh \u00bc h1 \u00f06\u00de where hh is the fish head deviation angle, h1 represents the swing angle of the joint 1, which can be obtained from (2). The longitudinal swing speed vh of the fish head is described as _yh, where yh is the fish head longitudinal position and defined as yh t\u00f0 \u00de \u00bc d y x1; t\u00f0 \u00de x1 \u00f07\u00de where d is the length of fish head and x1 is the length of the joint 1"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001298_tmag.2019.2942023-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001298_tmag.2019.2942023-Figure4-1.png",
        "caption": "Fig. 4. Magnetic field density distribution in the absence of the eccentricity of the rotor.",
        "texts": [
            " Unlike the analytical modeling in which the magnetic saturation effects of the stator and the rotor are neglected and they are infinitely permeable, the material DW540-50 is selected as the material of the stator and the rotor, in which the saturation flux density is around 1.46 T. The major parameters of active magnetic bearing, which is used for validation, are listed in Table I. A. Magnetic Flux Density Magnetic flux density distribution within the magnetic bearing plays a crucially important role in understanding the magnetic bearing and obtaining other significant characteristics. Fig. 4 shows the magnetic field distribution in the absence of the eccentricity of the rotor, which is calculated by the FEM software. The maximum value of the magnetic flux density is around 1.03 T, which is lower than the saturation flux density. The radial component of the magnetic flux density along the surface of the stator calculated by the analytical method is plotted, as shown in Fig. 5, which shows a good agreement with the FEM result with a maximum error of 2%. There are only some differences in the slots, in which the result of FEM is lower than the result of the analytical model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001237_j.mechmachtheory.2019.103652-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001237_j.mechmachtheory.2019.103652-Figure1-1.png",
        "caption": "Fig. 1. Translating cam system.",
        "texts": [
            " Norton [60] studied a similar problem for a double dwell cam system on a limited set of interpolating points without considering decimal truncation and comparing actual cam production techniques to traditional manufacturing approaches. The proposed study aims to verify if similar conclusions can be obtained also in the case of a single dwell cam system, testing new types of interpolation on a wide set of interpolating points and considering decimal truncation. A translating cam system with knife edge follower is considered, as shown in Fig. 1. Neglecting dynamical phenomena, the cam translates and impresses a desired motion y(t) to the follower. Due to the presence of: inertial properties m, a finite stiffness k and dissipation source c, the motion of the follower is described by the temporal function x(t), that, usually, it differs from y(t). The cam is supposed to translate perpendicularly to the motion of the follower and with a constant speed, thus there is a direct linear correlation between the physical shape of the cam and the profile of motion y(t)",
            "02% of the stroke and after 800 points, the error is closed to an asymptote, which value depends on the truncation decimals, as described in the empirical expression in (19). As mentioned, when the number of points and the truncation level are sufficiently high, the interpolation does not play an important role. With a very high number of points (that means a high truncation level) the simple linear interpolation is an interesting choice, because it allows the possibility to directly implement linear algebra. The approach proposed in Section 3 is applied to the problem shown in Fig. 1: a translating cam profile with constant speed coupled to a translating knife edge follower through a compliant transmission characterized by a finite stiffness [61] and a source of dissipation [62]. The results will drive the study through observations of simple dynamical phenomena to understand the effects of discretization, truncation and interpolation. This approach is characterized by a single degree of freedom and it is valid in different practical cases [5,53]. When the follower is very elastic, different vibrational modes can emerge, thus the proposed model is not valid to describe the dynamics of the system"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002746_j.triboint.2021.106920-Figure22-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002746_j.triboint.2021.106920-Figure22-1.png",
        "caption": "Fig. 22. Grid-shaped texture.",
        "texts": [
            " Combined with the flowrate calculation results of the two strokes, the average leakage can be obtained (see Fig. 21). The average leakage was 1.62594mL/h without texture and was 2.35495mL/h with grooves of 1\u03bcm depth and a spacing of 0.04mm and increased by 44.84% because the pressure difference created more leakage in the spiral groove. Therefore, the spiral groove texture degraded the seal performance of the O-ring under current operating conditions. Fig. 23 shows the virtual heights that were calculated for different groove depths and groove spacings (see Fig. 22). The variation trend of film thickness was consistent with the previous two textures, but the absolute value was similar to the first texture and higher than the second texture because, although the grid-shaped texture had connected pressure relief channels, the nodes where the grooves intersected produced a throttling effect. Fig. 24 shows the average pressure results that were calculated for different groove depths and groove spacings. The results for the gridshaped texture were similar to those of the first texture, the fluid film pressure increased and the asperities contact pressure decreased as the groove depth increased or the groove spacing decreased"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002217_j.neucom.2020.01.026-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002217_j.neucom.2020.01.026-Figure5-1.png",
        "caption": "Fig. 5. Quanser two-link flexible-joint rigid-link robot.",
        "texts": [
            " The results are compared with LQR in order to give the reader a notion of the difficulty of controlling our highly flexible and nonlinear system; we do not suggest LQR controls are used in practice for such a system. For fair comparison to an existing nonlinear control method we simply compare to adaptive backstepping using the standard e -modification robust weight update, which does as well as any controller in the literature that does not require precise knowledge of the model (to our knowledge). 4.1. Simulations of flexible-joint robot Our experiment comes from Quanser ( Fig. 5 ) and they provide specifications: link lengths l 1 = 0 . 34 , l 2 = 0 . 26 m, distance from the first joint to the first link\u2019s center of mass c 1 = 0 . 16 m, distance from the second joint to the second link\u2019s center of mass c 2 = 0 . 055 m, link masses m 1 = 1 . 5 , m 2 = 0 . 87 kg, link inertias I 1 = 0 . 0392 , I 2 = 0 . 0081 kg m 2 (where I 2 includes motor 2), and joint stiffness (spring constants) K 1 = 9 Nm/rad and K 2 = 4 N m/rad. Spong\u2019s model [30] with an assumption of large gear ratio gives M (\u03b82 ) \u0308\u03b8 = \u2212 C ( \u03b8, \u02d9 \u03b8) \u0307 \u03b8 \u2212 D \u03b8 \u02d9 \u03b8 \u2212 K ( \u03b8 \u2212 \u03c6) , (42) J \u0308\u03c6 = \u2212 D \u03c6 \u02d9 \u03c6 \u2212 K ( \u03c6 \u2212 \u03b8) + u , (43) where \u03b8 \u2208 2 x 1 contains the link angles, \u03c6 \u2208 2 x 1 contains rotor angles after gear reduction, M \u2208 2 x 2 is the inertia matrix, C \u2208 2 x 2 contains Coriolis and centripetal terms, J = diag (0 "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002755_1077546321998559-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002755_1077546321998559-Figure5-1.png",
        "caption": "Figure 5. Time-varying load of rolling element one under constant load 75 Nm (radial force 562.95 N).",
        "texts": [
            " An example of spur gear is used to investigate the effects of the time-varying load obtained by the dynamic model of the gear-bearing system on the bearing load distribution and Parameter Value Inner diameter (mm) 30 Outer diameter (mm) 90 Ball diameter (mm) 17 Number of balls 6 Damping of bearing (Ns/m) 2000 fatigue life. Gear parameters and bearing parameters are listed in Tables 1 and 2, respectively. The procedure is described in the flowchart displayed in Figure 4. First, the bearing static model was used to analyze the internal load distribution. As shown in Figure 5, the rolling element load was obtained under the constant external load 75 Nm, ignoring the bearing clearance. The rolling element load varies with a period of Tb related to the shaft frequency (see equation (1)). The maximum value is 416 N. Figure 6 displays the rolling element load, obtained from the dynamic model of the gear-bearing system, varying with a period of Tb and Tm (meshing frequency period). The maximum value of the rolling element load is 714.6 N. This maximum value increases by 71"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001955_j.mechmachtheory.2020.104090-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001955_j.mechmachtheory.2020.104090-Figure3-1.png",
        "caption": "Fig. 3. Constraint surfaces in the form of six spheres in accordance with Eq. (5) , for the FKP of the SPM with a non-planar FP and planar MP corresponding to the data listed in Table 1 .",
        "texts": [
            " The loop-closure equations are formulated by locating the vertices p i of the MP in the fixed frame of reference, in two different ways \u2013 directly, in the LHS and as a part of the MP, in the RHS (see Fig. 2 a and Section 2.1 ): b i + l i = p i = p c + R t i , i = 1 , . . . , 6 . (4) Once the input to the i thleg, given by l i = \u221a l i \u00b7 l i , is specified, the point p i is constrained to move on a sphere (denoted by S i ) of radius l i , centred at b i because of the spherical joint located at that point, as seen in Fig. 3 . Equations of these spheres can be derived from Eq. (4) as: S i : ( p c + R t i \u2212 b i ) \u00b7 ( p c + R t i \u2212 b i ) \u2212 l 2 i = 0 , i = 1 , . . . , 6 ; (5) \u21d2 S i := p c \u00b7 p c + 2 t i \u00b7 ( R p c ) \u2212 2 b i \u00b7 ( R t i ) \u2212 2 b i \u00b7 p c + t i \u00b7 t i + b i \u00b7 b i \u2212 l 2 i = 0 . (6) It may be seen from Eq. (6) that all the six spheres pass through the point p c . In terms of geometry, therefore, the FKP is equivalent of finding the condition for the concurrence of six spheres at a point in space , as seen in Fig. 3 . It may be noted that this condition is independent of the architecture of the SPM under consideration , as Eqs. (4 , 5, 6) hold good for any geometric form of the FP and the MP . In the following these equations are simplified to an equivalent form which facilitates their solution. Eq. (6) are six simultaneous non-linear equations in { p x , p y , p z }, as well as { c 1 , c 2 , c 3 }. However, since the only nonlinear term in the variables { p x , p y , p z } in these equations is p c \u00b7 p c = p 2 x + p 2 y + p 2 z , which appears identically in all the six equations, the same can be eliminated from all but one equation through the following steps",
            " In the following, all the linear dimensions of various SPMs are considered to be unit-less and are presented here only for the purpose of illustrating the FKP of the discussed architectures numerically. The unit of all angular quantities is radians. The architecture parameters for the present example have been adopted from [7] . Certain modifications have been performed to make the MP planar. Corresponding to this data, the concurrence of the six spheres, S 1 , . . . , S 6 , has been shown in Fig. 3 . The subsequent reduction of these to the equivalent geometric problem of concurrence of the planes P 21 , . . . , P 61 with the sphere S 1 has been depicted in Fig. 4 . Further, it is shown in Fig. 5 that the point p 1 can be determined uniquely if there exist three linearly independent planes passing through it. In the present example, the planes P 21 , P 31 , P 41 serve this purpose, as shown in Fig. 5 . For the architecture parameters and the input leg lengths presented in Table 1 the 64-degree FKU is computed as per Section 3"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002121_s12555-019-0625-0-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002121_s12555-019-0625-0-Figure9-1.png",
        "caption": "Fig. 9. Hybrid-sided varying domain optimization model.",
        "texts": [
            " When fuzzy inequalities and equality relationships exist simultaneously, the following hybrid-sided varying domain optimization model is formulated. max(\u03b4 \u2212\u03bb\u03b7) s.t. 0\u2264 fk \u2032(z)\u2264 \u03b2k\u2212\u03b4\u03b2k,k \u2208 I1, \u03b4\u03b2 j\u2212\u03b2 j \u2264 f j \u2032(z)\u2264 \u03b2 j\u2212\u03b4\u03b2 j, j \u2208 I2, \u03b4 = 1\u2212| fL \u2032(z)|,0\u2264 \u03b4 \u2264 1, \u03b2s\u2212\u03b2t \u2264 \u03b7 ,\u03b2t \u2212\u03b2L \u2264 \u03b7 , 0\u2264 \u03b2k \u2264 1,0\u2264 \u03b2 j \u2264 1,\u03b2L = 1, s, t,L \u2208 {1, . . . ,m} ,s 6= t 6= L, j = 1,2, . . .m1, j 6= L, k = 1,2, . . .m2, j 6= L, z \u2208 G(z), (35) where I1 is fuzzy inequality set, and I2 is fuzzy equality set. m1 is the number of fuzzy inequalities, and m2 is the number of fuzzy equalities. This case is shown in Fig. 9. 3.2.2 Fuzzy multiobjective cooperative path planning with different priorities via GVD Combining the multiobjective with different priorities and single-side varying domain optimization method, the multiobjective cooperative path planning model via GVD is built as max(\u03b4 \u2212\u03bb\u03b7) s.t. f\u0303 i 1 = 1\u2212 f\u0304 i 1, f\u0303 i 2 = 1\u2212 f\u0304 i 2, f\u0303 i 3 = 1\u2212 f\u0304 i 3, 0\u2264 f\u0303 i 1 \u2264 \u03b21\u2212\u03b4\u03b21, \u03b4 = 1\u2212| f\u0303 i 2|= 1\u2212| f\u0303 i 3|,0\u2264 \u03b4 \u2264 1, \u03b21\u2212\u03b22 \u2264 \u03b7 ,\u03b21\u2212\u03b23 \u2264 \u03b7 , 0\u2264 \u03b21 \u2264 1,\u03b22 = \u03b23 = 1. (36) 3.3. Distributed cooperative path planning via DMPC-GVD In order to obtain a feasible path, all the constraints in predictive horizon must be satisfied"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001939_s00521-020-05344-1-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001939_s00521-020-05344-1-Figure1-1.png",
        "caption": "Fig. 1 d\u2013q axis model of the PMSM",
        "texts": [
            " The conclusion is presented in Sec. 5. 2 Mathematical model and preliminaries The mathematical model of PMSMs based on d q rotating frame is given by [25] dH dt \u00bc x J dx dt \u00bc 3 2 np\u00f0Ld Lq\u00deidiq \u00fe 3 2 npUiq TL Bx Lq diq dt \u00bc Riq npLdxid npUx\u00fe uq Ld did dt \u00bc Rid \u00fe npLqxiq \u00fe ud 8 >> > > > > < > >> > >> >: \u00f01\u00de where H, x, iq and id stand for state variables, which represent the rotor angle, angular velocity, q-axis stator current and d-axis stator current, respectively. The system parameters are listed in Table 1. Figure 1 shows the d q axis model of the PMSM, where xr \u00bc npx and hr \u00bc npH. For simplicity, the following notations are defined: x1 \u00bc H; x2 \u00bc x; x3 \u00bc iq; x4 \u00bc id. Then, we have dx1 dt \u00bc x2 dx2 dt \u00bc a1 J x3 \u00fe a2 J x3x4 B J x2 TL J dx3 dt \u00bc b1x3 \u00fe b2x2x4 \u00fe b3x2 \u00fe b4uq dx4 dt \u00bc c1x4 \u00fe c2x2x3 \u00fe c3ud y \u00bc x1 8 > >> > > > >> < > >> > > > > >: \u00f02\u00de where uq and ud are the inputs of the system, y is the output and a1 \u00bc 3 2 npU; a2 \u00bc 3 2 np\u00f0Ld Lq\u00de; b1 \u00bc R Lq ; b2 \u00bc npLd Lq ; b3 \u00bc npU Lq ; b4 \u00bc 1 Lq ; c1 \u00bc R Ld ; c2 \u00bc npLq Ld ; c3 \u00bc 1 Ld \u00f03\u00de Fractional calculus can be defined in different ways, such as Gru\u0308nwald\u2013Letnikov, Riemann\u2013Liouville and Caputo"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000978_iemdc.2019.8785203-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000978_iemdc.2019.8785203-Figure5-1.png",
        "caption": "Fig. 5. Test-rig setup",
        "texts": [
            " Mechanical calibration for absolute strain sensitivity determination in a rotary application is beyond the scope of this study; relative strain measurement is deemed sufficient to enable understanding of diagnostic related information. Calibrating the sensor behaviour under exclusively thermal excitation conditions however can facilitate the differentiation of in-service signatures arising from thermal excitation from those produced by mechanical excitation, which will inherently be simultaneously registered by the insitu sensing head. Fig. 5 shows the schematic diagram of the test-rig system used in this study. The 0.55kW IM was driven by a commercial drive (Parker SDD890) operating in constant V/f control mode. For loading purposes, the IM was coupled to 0.75kW DC permanent-magnet machine whose armature current was controlled to achieve a desired operating point. The FBG sensors are interrogated by a SmatFibres SmartScan04 platform and processed using its proprietary LabView based SmartScan routine. The reflected Bragg wavelength from the installed FBG sensor was acquired at a frequency of 5 kHz in the test"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001352_aer.2019.130-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001352_aer.2019.130-Figure1-1.png",
        "caption": "Figure 1. Relative motion between missile and target.",
        "texts": [
            " 3\u00a9 The second-order sliding mode guidance law is designed by selecting the non-singular terminal sliding mode surface, which converges fast and can restrain chattering phenomenon effectively. 4\u00a9 The guidance law proposed in the paper can be used to attack maneuvering targets in 3D space with impact-angle constraints and has much value for engineering application. 2.0 3D SPACE TERMINAL GUIDANCE SYSTEM MODEL Considering the case of a missile attacking a ground maneuvering target in 3D space, the relative motion relationship between the missile and the target is shown in Fig. 1. In Fig. 1, Oxyz is the ground inertial coordinate system; M and T stand for the missile and the target respectively; r stands for the line of sight between missile and target; q\u03b5 and q\u03b2 stand for line-of-sight dip angle and line-of-sight drift angle, respectively; and the direction is defined as follows: q\u03b5 is positive when r is above the horizontal plane Oxz; q\u03b2 is positive when the ox axis rotates anticlockwise on the projection of r on the plane Oxz. vm is the velocity of missile, \u03b8m and \u03d5m stand for ballistic inclination angle and ballistic deflection angle. The direction is defined as follows: \u03b8m is positive when vm is above the horizontal plane; \u03d5m is positive when the ox axis rotates anticlockwise on the projection of vm on the plane Oxz. vt is the velocity of target. \u03b8t and \u03d5t respectively stands for course angle in pitching direction and in horizontal direction. Definition of their direction is the same as that of ballistic inclination angle and ballistic deflection angle. According to Fig. 1, the relative motion equation of missile target in 3D space is: \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 r\u0307 = vt ( cos \u03b8t cos q\u03b5 cos ( q\u03b2 \u2212 \u03d5t )+ sin \u03b8t sin q\u03b5 ) \u2212 vm ( cos \u03b8m cos q\u03b5 cos ( q\u03b2 \u2212 \u03d5m )+ sin \u03b8m sin q\u03b5 ) rq\u0307\u03b5 = \u2212vt ( cos \u03b8t sin q\u03b5 cos ( q\u03b2 \u2212 \u03d5t )\u2212 sin \u03b8t cos q\u03b5 ) + vm ( cos \u03b8m sin q\u03b5 cos ( q\u03b2 \u2212 \u03d5m )\u2212 sin \u03b8m cos q\u03b5 ) rq\u0307\u03b2 cos q\u03b5 = \u2212vt cos \u03b8t sin ( q\u03b2 \u2212 \u03d5t )+ vm cos \u03b8m sin ( q\u03b2 \u2212 \u03d5m ) . . . (1) The dynamic equation of the missile is: \u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 v\u0307m = ax \u03b8\u0307m = ay vm \u03d5\u0307m = \u2212 az vm cos \u03b8m . . . (2) Where, ax, ay and az are the three components of missile acceleration in velocity direction, normal velocity direction and lateral velocity direction"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002799_tie.2021.3063991-Figure24-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002799_tie.2021.3063991-Figure24-1.png",
        "caption": "Fig. 24. Exploded view of main parts of the proposed machine.",
        "texts": [
            "ieee.org/publications_standards/publications/rights/index.html for more information. VI. EXPERIMENTAL VERIFICATION In order to validate the above theoretical analysis and simulation results, a prototype machine is manufactured and tested. Fig. 23 shows the main components of prototype machine. The inner stator with armature winding, FW and PM is illustrated in Fig. 23(a), and the rotor structure is shown in Fig. 23(b). The prototype machine is fabricated according to the parameters in Table III. Fig. 24 depicts the exploded view of the main parts of the proposed machine. It contains stator, windings, rotor, housing, bearings, output shaft, and fixed shaft. The main parts and assembly process are similar to the conventional PM machine. The tested and simulated open-circuit back-EMF waveforms at 2.5 A/mm 2 , 0 A/mm 2 , -2.5 A/mm 2 FW current are plotted in Fig. 25, where the rotation speed is 120 r/min. It is shown that the back-EMF waveforms are quite sinusoidal and the results match well. Thus, regulation of back-EMF by adjusting FW currents of the prototype machine is verified"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001312_tie.2019.2952780-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001312_tie.2019.2952780-Figure12-1.png",
        "caption": "Fig. 12. Experimental setup: (a) For detecting the thrust of the electromechanical actuator with the spindle rotational speed, (b) For detecting the clamping force of the chuck mounted on the electromechanical actuator using a clamping force gauge.",
        "texts": [
            " 1) Test for the thrust produced by the drawbar during the clamping: To analyse the thrust produced by the drawbar during the clamping mode, the drive motor shaft is coupled with the motor side shaft of the electromechanical actuator using a pulley-belt mechanism as shown in Fig. 9 and the motor is run at the low speed winding mode. The pulley-belt mechanism used for the test bench has a pulley transmission ratio of 1.5. The clutch system highlighted in Fig. 9, connected to the stationary side as mentioned in section II, to restrict the spindle system of the electromechanical actuator from rotating. As shown in Fig. 12 (a), a load cell (Kistler strain gauge meter4703) is attached at the end of the drawbar, to measure the thrust produced by the drawbar by transforming the rotary motion of the drive motor, transmitted through the pulleybelt mechanism into a linear motion. A Yaskawa inverter (A1000) is used to perform the speed control of the drive motor. Drive motor torque command is varied between 20 Nm to 70 Nm, and the produced torque by motor along with the thrust generated by the drawbar are measured repeatedly for each torque command value of the drive motor which are listed in Table IV",
            " With this, in (8), the only unknown was the friction cofficient which is found by putting the values of other parameters in (8). To ensure accuracy, a repetitive experimental data for the thrust and torque are used while calculating the friction coefficient. The difference between the analytical and the experimental data arises beacuse of the manufacturing tolerances associated with the different stages of the actuator, during the power transmission from the motor side to the drawbar side. 2) Test for chuck clamping force by the chuck jaw: To detect the chucking clamping force, as shown in Fig. 12(b), a three jaw power chuck is added at the end of the drawbar and a chuck force gauge (Schunk-GFX-270) is place in between. The aim of this test is to measure the chucking force at the tip of the jaws while holding a workpiece for machining operation. It is essential for a lathe system to ensure a high and equal amount of clamping force on each jaw while holding a workpiece, because an unbalanced clamping force can cause defective end products. Furthermore, unbalanced holding or reduction of clamping force during machining may results in fatal accidents"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001817_lra.2020.3013882-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001817_lra.2020.3013882-Figure4-1.png",
        "caption": "Fig. 4. Experimental environment. (a) Image showing the experimental environment around TCPF, and (b) overview of the experimental environment.",
        "texts": [
            " Then, we wound the twisted thread onto a mandrel with a diameter of 1.6 mm. To fix the thread in a coil shape, we heated the thread at 180 \u00b0C for 10min by an automatic oven (DOV-450, AS-ONE). We first evaluate the performance of the developed COF sensor when driving the TCPF actuator and then model the developed COF sensor. As this letter focuses on investigating whether the COF sensor is available when driving the TCPF actuator and changing the load, a practical model is established in this section. The experimental environment around the TCPF is shown in Fig. 4. In the experiment, we attached a COF in parallel with the TCPF actuator. Because we supposed that the actuator unit with the developed COF sensor is driven in an air-cooling system, a cooling fan was used in the subsequent experiments. To prevent the TCPF actuator from melting by over-heating, its temperature was monitored using a thermography camera (OPTX180LTF05T090, Optris). For evaluation and modeling, the displacement measured by a laser range sensor (IL-300, KEYENCE) was used as the ground truth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001218_s42417-019-00182-5-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001218_s42417-019-00182-5-Figure4-1.png",
        "caption": "Fig. 4 Geometry of the milling edge of the end milling cutter",
        "texts": [
            " While the s , z and h represent the length of the milling edge, the axial cutting depth and the cutting thickness of the end milling cutter, respectively. The Ktc,Krc and Kac are tangential, radial and axial the cutting force coefficients, respectively. While the Kte,Kre and Kae are the tangential, radial and axial edge force coefficients, respectively. (1) \u23a7\u23aa\u23a8\u23aa\u23a9 dFt = Ktc \u22c5 h \u22c5 dz + Kte \u22c5 ds dFr = Krc \u22c5 h \u22c5 dz + Kre \u22c5 ds dFa = Kac \u22c5 h \u22c5 dz + Kae \u22c5 ds , 1 3 We set the number of milling cutter teeth as N, radius as R, helix angle as i , axial cutting depth as ap , and radial cutting depth as ae . The geometry of the milling edge is shown in Fig.\u00a04. If the tooth distribution on the milling cutter is uniform, the angle between the teeth of the milling cutter is p = 2 \u2215N . Assuming that the angular displacement at the first microelement of the first milling edge is 10 , the instantaneous radial contact angle at the jth milling edge and lth microelement is given by: During the milling process, the milling thickness changes with the milling edge, the relationship between the milling thickness and the turning angle of the cutter teeth is approximated as: where ft is the feed per tooth"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001999_ecce44975.2020.9235600-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001999_ecce44975.2020.9235600-Figure4-1.png",
        "caption": "Fig. 4. d-axis magnetic flux when applying magnetizing current.",
        "texts": [
            " On the other hand, the proposed VFMM doesn\u2019t use positive magnetization range since this model can generate large reluctance torque without large magnetic flux obtained from PMs. In addition, it is necessary that operating points don\u2019t exceed the knee point when applying load current. If operating points of VPMs exceed the knee points, unintentional demagnetization occur and the average torque is dramatically reduced. Furthermore, operation points are on the initial magnetization curve in case that magnetization ratio is 0% in the proposed VFMM as shown in Fig. 3(b). Fig. 4 shows d-axis magnetic flux \u03a8d when applying magnetizing current pulse in the conventional and the proposed VFMMs respectively. Both models employ deltatype PM arrangement which combines a first layer CPM with second layer V-type CPMs. The first layer CPM can restrain the diamagnetic field from second layer CPMs which might cause unintentional demagnetization of VPMs [4]. Moreover, in the conventional VFMM indicated in Fig. 4(a), most of the d-axis magnetic flux is concentrated to VPMs because the flux barriers which is installed between poles can be sufficiently extended. As a result, \u03a8vpm that contributes to the magnetization of VPMs becomes large. Therefore, the conventional VFMM can use magnetization ratio of VPMs from -100% to +100%. On the other hand, in the proposed VFMM shown in Fig. 4(b), distance between first and second layer CPMs is increased in order to enhance q-axis magnetic flux \u03a8q while load current is applied. Therefore, the proposed VFMM has a potential to obtain large reluctance torque, but extended flux barriers become smaller due to wide distance between first and second layer CPMs. Consequently, smaller extended flux barrier causes increase in leakage flux \u03a8l that passes extended flux barriers as shown in Fig. 4(b). In other words, \u03a8vpm that contributes to the magnetization of VPMs becomes small. Accordingly, the proposed VFMM doesn\u2019t use positive magnetized range of VPMs as shown in Fig. 3(b). \u03a8q VPMVPM CPM \u03a8q Reduction (Saturation) VPMVPM CPM (a) The conventional VFMM Reinforcement ribs due to large q-axis magnetic flux \u03a8q. Therefore, the proposed VFMM can achieve the target torque of 163 Nm without positively magnetized VPMs. In addition to that, if the reluctance torque is dominant, torque density can become large easily in traction applications because the reluctance torque is proportional to the square of load current"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001905_j.mechmachtheory.2020.104095-Figure13-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001905_j.mechmachtheory.2020.104095-Figure13-1.png",
        "caption": "Fig. 13. Vibrating around axis y p .",
        "texts": [
            " To facilitate laser alignment at the target, a piece of black tape (about 20mm \u00d7 20mm in dimension) is attached to the end of the cantilever beam. The target of the laser vibrometer is stuck at the black tape. Before testing, we measured the background noise. The amplitude of background noise is about 0.8\u03bcm, and the frequency of background noise is about 30-40Hz. Then, two typical vibration modes are completed to test the accuracy of the dynamic model. The typical vibration modes are vibrating around axis y p and vibrating around axis x p . Case I: vibrating around axis y p . Fig. 13 shows the vibration mode of case I. The dotted arrow means the vibration direction of P 0 . The solid arrow means the vibration direction of P 1 . In case I, only sinusoidal drive signals are applied to the piezoelectric ceramic plates of kinematic 1. The driving signals of the two piezoelectric ceramic plates on kinematic chain 1 have the same amplitude, frequency, and phase. Numerical results and experimental results are shown in Fig. 14 . According to the datasheet of the piezoelectric ceramic plate, the output force of the piezoelectric ceramic plate is linear with voltage"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003766_0094-114x(95)00069-b-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003766_0094-114x(95)00069-b-Figure1-1.png",
        "caption": "Fig. 1. Geometric parameters for transformation of twists. {f} and {g} are reference frames fixed to the world. {m} and {n} are reference frames fixed to the moving rigid body. The vectors p0 link the origin of frame {i} to the origin of frame {j}.",
        "texts": [
            " Choice of six-vector This text consistently uses the following convention for a twist six-vector t: to is the angular velocity three-vector of the rigid body; v is the velocity three-vector of the reference point on the rigid body. (The alternative choice, t = (VT~T) T, is also in widespread use.) If it is necessary to explicitly indicate that the components of t are expressed in a reference frame {i} then a leading subscript \"i\" is added: it. If it is necessary to explicitly mention the velocity reference point on the rigid body--for example, the origin of the frame {j }--a trailing superscript '7\" is added: tL 2.2. Choice of a reference frame The left-hand side of Fig. 1 shows two reference frames {f} and {g} fixed to the world, and a reference frame {m } fixed to the moving rigid body. The origin of {m} serves as the velocity reference point. Let the twist of the rigid body with respect to the frames {f} and {g} be given by the six-vectors \\ i v = ) ' \\ =vm ) \" (The trailing superscript \"m\" will be omitted for the rest of this paragraph since the velocity reference point remains unchanged.) Physically, the angular velocity vector to is the same for both {f} and {g}, as well as the velocity vector v",
            " A wrench is a six-vector containing a linear force three-vector and a torque three-vector. See for example [12] for an extensive review on screw theory.) ~P does not change the physical three-vectors of the twist, but just projects them onto another reference frame; this projection changes the numerical values: ( ~R 03 \u00d7 3) s t =~P~t = \\03\u00d73 ~R _st\" The inverse transformation is simply (~P) ' =~P. Remark that ~P does not depend on the choice of velocity reference point. (4) 2.3. Choice of velocity reference point The right-hand side of Fig. 1 shows two reference points, the origins of the frames {m } and {n }, fixed to the rigid body, and a reference frame {f} fixed to the world. Let the twist of the rigid body with the origin of {n } as velocity reference point be given by the six-vector <) jt n = fv n (The leading subscript ' ~ ' will be omitted for the rest of this paragraph since all motion parameters are expressed with respect to the same frame {f}.) The physical angular velocity e~ does not change under the transformation of the velocity reference point from the origin of {n } to the origin of {m }, but the physical translational velocity vector v does: 138 H"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001948_icra40945.2020.9196780-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001948_icra40945.2020.9196780-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagrams and prototypes of continuum manipulator ((a)(b)) and soft gripper((c)-(e))",
        "texts": [
            " Numerous designs of underwater continuum robots were presented in [12], [13] and [14]. The water will be easier to obtain and carry than air and more environmentally friendly than hydraulic oil. Although the underwater manipulators mainly use oil hydraulic cylinders and motors, the load capacity, reliability and maintainability of the conventional oil hydraulic driving machinery are reduced due to their large weight and susceptibility to the underwater environment [15]. Therefore, the underwater multi-segment continuum manipulator driven by McKibben WHAM was developed. Fig. 1(a) shows the 3D-CAD drawing of the multi-segment manipulator. The manipulator consists of 3 segments. One segment has a soft gripper (model in Fig. 1(c)) and three vertical WHAMs. The proposed soft gripper has 3 linked planar 978-1-7281-7395-5/20/$31.00 \u00a92020 IEEE 2946 Authorized licensed use limited to: Heriot-Watt University. Downloaded on September 23,2020 at 14:38:54 UTC from IEEE Xplore. Restrictions apply. linkages driven by 3 transverse WHAMs. The depressurized length of the transverse WHAMs is 60 mm, while the vertical WHAM\u2019s is 80 mm. The gripper fingers was made by rubber slab, which can adapt to the gripping objects under the press deformation. Fig. 1(b) presents the final manipulator. When the transverse WHAMs of 3 segments are pressurised, the grasping diameter of the soft grippers will decrease, as the red circles shows in Fig. 1(d) and (e). When the soft gripper of one segment is driven, the 3 vertical WHAMs of this segment will also bend. The transverse and vertical WHAMs will influence each other while pressurised and depressurised. So the compliance of the continuum robot is an essential characteristics for the manipulator in this paper. Especially when the manipulator bends, it is designed to grasp objects meanwhile. The rigid structure can not meet the desired function. And the compliance of the WHAM can perfectly realise both the flexibility and certain hardness",
            " After the artificial muscles are pressurized and contracted, the grasping mechanisms are driven to perform a zooming motion to grasp objects. Based on the graphical relationship of the planar linkage 2947 Authorized licensed use limited to: Heriot-Watt University. Downloaded on September 23,2020 at 14:38:54 UTC from IEEE Xplore. Restrictions apply. mechanism, the relationship between the contraction length of the 3 transverse WHAMs and the gripping diameter can be calculated, as shown in Fig. 6. The diameter of the depressurized soft gripper (inner circle in the right Fig. 1) is 120 mm. When L > 21.1mm, the gripping diameter will be negative, which means that the 3 gripper fingers will interfere with each other in space and may be destroyed. According to the experimental data of the top plot in Fig. 3, the contraction length is less than 13 mm, which confirm the safety of the motion of soft grippers. Besides, the curve in Fig. 6 is almost a straight line, which benefits the linear control of gripping diameter. When the transverse grippers are not actuated, the continuum manipulator conforms to the piecewise constant curvature hypothesis; When the transverse grippers are actuated, the distribution diameter of longitudinal grippers changes"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure41.4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure41.4-1.png",
        "caption": "Fig. 41.4 Retractable roof in expanded (a) and retracted (b) state",
        "texts": [
            " After completion of functional block \u00abUAV battery replacement\u00bb, the station switches to mode \u00abBattery storage\u00bb until the next replacement mission. Two-way data transfer module enables data exchange at distance up to 50 kmwith operational frequency grids of 2.3/2.4/2.5 and 5.2\u20135.8 GHz. Maximum data transfer speed is 12Mbit/s. Video stream can be transferred with resolution of 1920\u00d7 1080p, frequency 30 frames/sec. and delay 125 ms. To obtain publicly available meteodata, a 4G modem is used. To implement functional blocks \u00abUAV landing\u00bb, \u00abUAV storage\u00bb and \u00abUAV takeoff\u00bb, a retractable roof was developed (see Fig. 41.4). The roof protects the vehicle from poor weather conditions during storage of the UAV, as well mechanical and electronic components of the station. The retractable roof consists of five polycarbonate rectangular sections, where subsequently solar panels should be placed to reduce base station energy consumption from the battery. Roof retraction is performed along the stainless steel rails with two step motors and cord drive. Themain advantages of the developed retractable roof are the following: compactness in retracted state, protection of UAV and bottom level mechanisms from poor conditions, possibility to mount solar panels on the roof, low energy consumption"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002066_01691864.2020.1854115-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002066_01691864.2020.1854115-Figure4-1.png",
        "caption": "Figure 4. The shoulder complex of the exoskeleton is comprised of three joints, one active joint, Joint 4, and two passive joints, Joint 2 and Joint 3. The ICR of the shoulder complex of the exoskeleton should coincide with that of the wearer in operation.",
        "texts": [
            " Among these four joints, GH joint is considered as a ball-and-socket joint, acting as the most important structure to determine the range of movement of the shoulder complex. Therefore, we first consider the three degrees of freedom of the articulations corresponding to abduction/adduction, flexion/extension, and horizontal abduction and adduction. The shoulder complex of the exoskeleton is designed to possess the above three Figure 2. The configuration of the active joint and the rotatable and telescopic mechanism of the exoskeleton\u2019s elbow joint. degrees of freedom as shown by Joint 2, 3, and 4 in Figure 4. Joint 4 corresponding to flexion/extension movement is an active joint in almost the same structure as the active elbow joint as well as its safety precautions. Hence, the description about the active Joint 4 is omitted here. Joint 2 and Joint 3, designed as passive joints, are structured as a simple revolute shaft and hinge. They are mainly used for following the abduction/adduction, and horizontal abduction and adduction of the wearer, respectively. These two joints are orthogonal to each Figure 5",
            " Thus, the shoulder complex of the exoskeleton gradually raises to its highest posture when the back support shafts rotate to the vertical posture that is parallel to the vertical prop as shown in Figure 6(a). Since the length of the back support shafts and their initial inclination angles can be elaborately calculated, the elevation/depression range of the exoskeleton\u2019s shoulder complex is just about 12mm (as x shown in Figure 6). In above design, there is an offset of 100mm between Joint 3 and the intersection point O, as shown in Figure 4, which results in a displacement of elevation/depression when the wearer undergoes abduction/adduction movement. Nonetheless, Figure 6(b) shows that the rotation of the back support shaft about Joint 1 causes a lateral deviation of approximate 100mm from Joint 3. This deviation would just compensate for the above displacement (100mm) between Joint 3 and the wearer\u2019s ICR. This design effectively utilizes the following ability of the passive joint to guarantee that the ICR of the exoskeleton\u2019s shoulder complex coincides with the ICR of the wearer\u2019s shoulder"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002352_j.mechmachtheory.2020.103945-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002352_j.mechmachtheory.2020.103945-Figure9-1.png",
        "caption": "Fig. 9. An umbrella\u2013shaped mechanism: (a) folded configuration, and (b) deployed configuration.",
        "texts": [
            " Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx An umbrella-shaped deployable mechanism can be constructed by connecting a serial of identical deployable units through some assembly methods[34]. For example, by sharing a FB and connecting every two units through an R pair, six \u201c \u02dc R l1 \u22a5 \u2194 R l2 \u02dc R c1 / \u0303 R c2 \u02dc R u 1 \u22a5 \u2194 R u 2 \u22a5 \u0303 R u 3 \u201d deployable units can construct an umbrella-shaped deployable mechanism, whose folded configuration and deployed configuration are shown in Fig. 9 (a) and (b). From Fig. 9 , an umbrella-shaped deployable mechanism can hold the plane-symmetry and centrosymmetry from the folded configuration to the deployed configuration, so the mechanism has the same kinematics as its linkage units. Here, a general method for kinematics analysis of all kinds of two-layer and two-loop deployable linkage units is developed. 3.1. Displacement analysis of two-layer and two-loop deployable linkages For convenience, in the external loop linkage of a deployable unit, the moving link always passing through the SP per- pendicularly is called the end link , and a pair whose axis is in the SP or which connects with the end link is called an end pair "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000983_j.compgeo.2019.04.023-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000983_j.compgeo.2019.04.023-Figure7-1.png",
        "caption": "Fig. 7. Configuration of a sphere on an incline.",
        "texts": [
            " Contact forces between spheres and the ground are denoted as F1, F2 and F3, respectively. Three spheres have the same material properties: density \u03c1 = 2.0 \u00d7 103 kg/m3, radius R = 1.0 m, Young\u2019s modulus E = 2 GPa, and Poisson\u2019s ratio = 0.2. The acceleration of gravity is given by \u221210 m/s2, and the time step is set as 0.02 s. The numerical results are summarised in Table 2. A good agreement between analytical solutions and numerical results is achieved. A sphere rests on an incline at an angle of 45\u00b0 from horizontal, as shown in Fig. 7. The sphere slides down under the action of gravitational force and the surface friction angle was varied from 0\u00b0, 10\u00b0 to 15\u00b0. The material density of the sphere is given by 2.0 \u00d7 103 kg/m3; the radius is 1.0 m; the elastic modulus is 200 GPa and the Poisson\u2019s ratio is 0.20. The model parameters include the time step of 0.01 s, the acceleration of gravity of \u221210 m/s2 and a zero rolling friction coefficient. In the case when the friction angle is 15\u00b0, the same numerical test using the classic DDA was conducted in [66]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002044_ccece47787.2020.9255787-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002044_ccece47787.2020.9255787-Figure1-1.png",
        "caption": "Fig. 1. Traditional permanent magnet machines [3], [4]. (a) Radial-flux PMSM. (b) Axial-flux PMSM. (c) SPMSM. (d) IPMSM.",
        "texts": [
            " So far, two main types of electric machines have been employed in the passenger car industry:  Induction motors (IMs)  Permanent magnet synchronous motors (PMSMs) Although it has drawbacks such as cost and complexity [2], the PMSM architecture is often preferred because of its high efficiency and power density [4]. Moreover, PMSMs can be designed with different configurations, classified by flux direction and magnet positioning [3] as follows:  Radial-flux PMSM o Internal rotor PMSM  Surface-mounted PMSM (SPMSM)  Interior permanent magnets synchronous machine (IPMSM) o External rotor PMSM  Axial-flux PMSM Figure 1 shows some examples of electric motors, highlighting the differences in their architecture. As mentioned in [6], [7], the IPMSM architecture is the most commonly employed one for electric traction motors in electrified vehicles due to its advantageous performance in efficiency and power density with respect to the other architectures, and it is therefore the subject of this paper. Even before electrification was a major trend in the automotive industry, guaranteeing comfort to the passengers has been one of the most important concerns for designers and engineers, and this has not changed in the meantime; one of the crucial aspects for powertrains that is related to this issue is therefore their noise and vibration production, as it can heavily influence riding comfort"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001357_j.oceaneng.2019.106812-Figure16-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001357_j.oceaneng.2019.106812-Figure16-1.png",
        "caption": "Fig. 16. Quarter FE model of rack and cylinder welded structure.",
        "texts": [
            " Ocean Engineering 196 (2020) 106812 time, and computational time of mechanical analysis can be reduced by 70% by earlier iterative substructure method and reduced by 82% by proposed iterative substructure method. In addition, the proposed iterative substructure method with parallel computation can reduce about 40% computational time comparing to earlier iterative substructure method. In order to accurately predict the welding deformation and validate with measured data, quarter FE model of rack and cylinder welded structure as shown in Fig. 16 was used because of geometrical symmetry and computation resource reduction. In Fig. 16, there are 61 multipass welding seam with different color belong to 12 layers in V type welding groove, and the welding passes located at the left and right of rack will be welded symmetrically and simultaneously. The total point and element are 122,859 and 114,350, respectively. The horizontal direction of points on line1 to line 2 and upright direction of points on line 3 to line 4 are all fixed to represent the symmetrical boundary condition, and rigid body motion is also constrained as mechanical boundary condition",
            " 19 shows the cross section shape of examined cylindrical leg structure after rack and cylinder welding, in which original and deformed shapes were marked with yellow and orange colors. It is also can be seen that same deformation trends comparing with experiment observation. Although the computed welding deformation has identical tendency comparing with experimental observation, it is still necessary to compare the magnitude of computed welding deformation and measured data. Computed Welding deformation of points on line 2 as indicated in Fig. 16 with shrinkage feature was compared to measurement with inward trend as shown in Fig. 20(a), in which good agreement can be observed. Similarly, Computed Welding deformation of points on line 4 as indicated in Fig. 16 with expansion feature was compared to measurement with outward trend as shown in Fig. 20(b), in which good agreement also can be observed. Noting to the comparison between measurement and computation, the measured welding distortions are located to the sections as shown in Fig. 5 with total length of 36,000 mm, and welding length of quarter FE model of rack and cylinder welded structure is 500 mm. Predicted welding distortion has a consistent feature, and can be used to approximately consider the welding distortion configuration of weld structure FE model with actual length"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001583_1077546320916628-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001583_1077546320916628-Figure4-1.png",
        "caption": "Figure 4. Measuring point position of pantograph vibration.",
        "texts": [
            " The experimental vibration data are obtained from four pantograph fault types which are healthy, horn crack fault, head bracket crack fault and composite fault. The fault components are shown in Figure 3. The length of head bracket crack is 30 mm and the length of horn crack is 15 mm. Vibration data are collected using a recorder at a sampling frequency of 5 kHz for different pantograph conditions. Three acceleration sensors were used to collect the panhead top pipe and front and rear carbon strip vibration signals, as shown in Figure 4 and Table 1. Tests are carried out with the speed of 60 km/h. Vertical vibration data of panhead top pipe under healthy, horn crack fault, head bracket crack fault, and composite fault are shown in Figure 5. To investigate the effect of parameters, according to equation (4), such as data lengthN , embedded dimensionm, Figure 5. Vertical vibration data of panhead top pipe under four conditions. Table 1. Measuring points. Number Measurement points 1 Vertical vibration of front carbon strip 2 Vertical vibration of rear carbon strip 3 Vertical vibration of panhead top pipe and time delay \u03c4 on PeEn calculation, relevant study has been conducted"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003947_1.2059196-FigureI-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003947_1.2059196-FigureI-1.png",
        "caption": "Fig. I. A schematic diagram of the batch slurry reactor; A, mechanical stirrer; B, platinum anode; C, anode compartment; D, five-necked flask; E, gas-dispersion tube; F, reference electrode (SCE); G, gold working electrode; H, stirrer blade.",
        "texts": [
            " I~ The primary advantages of a slurry electrode system over a fixed-bed Raney nickel powder reactor are: (/) the potential of all catalyst particles in contact with the gold target electrode is precisely known; thus traditional electrochemical methods of analyzing current-potential data can be used to obtain kinetic parameters (i.e., one does not need to solve a porous-electrode model 18 for the potential, current, and reaction-rate profiles in order to calculate kinetic parameters from current-voltage data) and (ii) the solution pH and the concentrations of reactant and product in the solution adjacent to all catalyst particles is known and equal to the bulk concentration. The electrochemical batch slurry reactor (shown in Fig. I) was similar to that employed by Baria and Hulburt 12 and consisted of a 400 ml multineeked glass flask containing a flat sheet puratronic gold strip (8.9 cm x 1.3 cm) working (target) electrode, a Luggin capillary probe connected to a remote saturated calomel reference electrode, a high purity platinum foil (I cm X i em) eounterelectrode (anode), and a mechanical stirrer. A medium-porosity sintered-glass disk separated the working and counterelectrode compartments and prevented catalyst particles from contacting the platinum foil anode",
            "2Downloaded on 2013-11-28 to IP and s in the glucose-reduction/H2-evolution model were determined, potential-sweep data and sorbitol production rates from batch slurry-reactor experiments were used as additional tests of the model A typical fit of the model with experimental currentvoltage data is shown in Fig. 8 for a solution temperature of 298 K, a glucose concentration of 0.2M, a solution pH of 7 with a suppoz~ting electrolyte composition of 0.5M Na2SQ + 0.025M KH2PQ + 0.025M Na2HPQ, and a catalyst loading of 0.051 g/cm 3. The match of the theory and data was achieved using FG = 1.37 \u2022 104 cm3/mol and S = 200 cm2/g. The dotted line in the figure is the H2 evolution current density-overpotential curve for Raney nickel powder in the absence of glucose, as computed from the VolmerHeyrovsk~ rate equation (Eq. 10) with # = 0.50 and values for Zo,v and Zo,Hy at pH 7 and T = 298 K from Table II. The results in Fig. 8 show that the H2 evolution current densities on Raney Ni are lowered (by a factor of three) when 0.20M glucose is present in the solution. Such an observation differs significantly from tradit ional direct-electrontransfer reductions with high-hydrogen-overpotential cathodes where the addition of an organic reactant causes the measured current to increase above background (see, for sample the current-voltage plots for the reduction of glucose to sorbitol on Zn(Hg)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002232_1350650119900401-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002232_1350650119900401-Figure1-1.png",
        "caption": "Figure 1. Loading and motion state of the bearing during start-up/shut-down. (a) Loading and motion state. (b) Angular velocity change of the inner race during start-up and shut-down. (c) Angular velocity change of the ball during shut-down. (d) Linear velocities and sliding velocities during shut-down.",
        "texts": [
            " Then, the mixed TEHL analysis is employed to obtain the changing rules of lubrication state and properties. The coupling method is implemented by the looping and iteration of bearing quasidynamic analysis and mixed TEHL analysis, which can provide theoretical supports for getting accurate bearing dynamic and lubrication properties, reasonable judgment of lubrication state and effective prediction of lubrication failure. Kinematic equations of ball bearing during start-up and shut-down The process of inner race start-up and shut-down is shown in Figure 1(b), during start-up, the time steps are set from t0 to tn, the initial inner race angular velocity is !2 t0\u00f0 \u00de \u00bc 0, then the angular velocity steps to a small value !2 t1\u00f0 \u00de and speeds up to !2 tn\u00f0 \u00de. During shut-down, it is an inverse process, the angular velocity of inner race slows down from !2 t1\u00f0 \u00de to !2 tn\u00f0 \u00de. The outer race remains static during start-up and shut-down, the ball bearing is loaded by the applied forces FX, FY, FZ in X-, Y- and Z-directions, respectively, and applied moments MY, MZ around Yand Z-directions",
            "z0j sin 2j Dm=\u00f02 cos 2j\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R 02 2 y22j q h ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R 02 2 a2j2 q ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0DW=2\u00de 2 a2j2 q \u00f02\u00de Ux2j2 \u00bc !2 !o0j cos 2j Dm= 2 cos 2j ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R022 y22j q ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R022 a2j2 q ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0DW=2\u00de 2 a2j2 q \u00f03\u00de Ux2j \u00bc Ux2j j \u00feUx2j2 2 \u00f04\u00de Sx2j \u00bc 2 Ux2jj Ux2j2 Ux2jj \u00feUx2j2 \u00f05\u00de Figure 1(c) and (d) shows the rotation and revolution velocities of the ball, the linear velocities of the ball and inner race, and the sliding velocity at the azimuth angle j \u00bc 0 during shut-down. As shown, the motion state of the ball and the inner race change apparently, and it can provide real running conditions for mixed TEHL analysis. In the analysis, the aero-engine main shaft ball bearing 476728NQ is taken as the research object, and the Newton\u2013 Raphson method and steepest descent method have been employed to conduct quasi-dynamic analysis"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002352_j.mechmachtheory.2020.103945-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002352_j.mechmachtheory.2020.103945-Figure12-1.png",
        "caption": "Fig. 12. Deployable unit \u201c \u02dc R l1 \u22a5 \u2194 R l2 \u02dc R c1 / \u0303 R c2 \u02dc R u 1 \u22a5 \u2194 R u 2 \u22a5 \u0303 R u 3 \u201d: (a) structural diagram, and (b) main dimensions.",
        "texts": [
            " (62) , (64) and (67) can satisfy the condition of Eq. (61) , the units with structural type \u201c{4R} [4R] (5R)\u201d developed in previous work [35] , and the units with structural types\u201c{4R} [3R] (6R)\u201d and \u201c{6R} [1R] (6R)\u201d developed in this work have simpler kinematics than the rest units. The new units with structure type \u201c{6R} [2R] (5R)\u201d have the strongest coupling between kinematic equations. 3.5. A typical numerical example For example, for the unit \u201c \u02dc R l1 \u22a5 \u2194 R l2 \u02dc R c1 / \u0303 R c2 \u02dc R u 1 \u22a5 \u2194 R u 2 \u22a5 \u0303 R u 3 \u201d shown in Fig. 12 (a), its main dimensions shown in Fig. 12 (b) are set as follows: l 0 = 56 mm, \u03b3 1 = 30 deg, l 1 = 50 mm, l 2 = 50 mm, h 1 = 20 mm, h 2 = 15 mm, l 3 = 50 mm, l 4 = 45 mm, l 5 = 20 mm, \u03b3 3 = 140 deg, l 6 = 60 mm. \u03b3 2 = 20 deg and l 7 = 12 mm. For such a unit, it can be easily found p = 2, q = 3 and t = 2. The homogeneous transformation of frame e with respect to frame 0 from chains 1 and 2 can be expressed as 0 e T ( \u03b8l1 , \u03b8l2 , \u03b8u 1 , \u03b8u 2 , \u03b8u 3 ) = 0 F T F 1 T ( \u03b8l1 ) 1 2 T ( \u03b8l2 ) 2 3 T ( \u03b8u 1 ) 3 4 T ( \u03b8u 2 ) 4 5 T 5 e T ( \u03b8u 3 ) = \u23a1 \u23a2 \u23a3 n xe o xe a xe r xe n ye o ye a ye r ye n ze o ze a ze r ze 0 0 0 1 \u23a4 \u23a5 \u23a6 (68) Please cite this article as: W"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002770_j.jmapro.2021.02.015-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002770_j.jmapro.2021.02.015-Figure3-1.png",
        "caption": "Fig. 3. The model of the SLMed circle structure with 67\u25e6 rotate scanning strategy: (a) layered by 67\u25e6 rotate; (b) overhead view.",
        "texts": [
            " Journal of Manufacturing Processes 64 (2021) 907\u2013915 be seen in Table 2, as the \u03b1n increases from 0\u25e6 to 45\u25e6 and 90\u25e6 to 135\u25e6, the Rne and Dnw decreases, while the Rni increases; as the \u03b1n increases from 45\u25e6 to 95\u25e6 and 135\u25e6 to 180\u25e6, the Rne and Dnw increases, while the Rni decreases. When the \u03b1n is 0\u25e6, 90\u25e6 and 180\u25e6, the maximum Rne, Dnw and minimum Rni can be obtained. When the \u03b1n is 45\u25e6 and 135\u25e6, the minimum Rne, Dnw and maximum Rni can be obtained. Therefore, the homogeneity of the dimension distribution is poor. Similar to the 90\u25e6 rotate scanning strategy, the model of the SLMed circle structure with 67\u25e6 rotate scanning strategy can be established, as shown in Fig. 3. Differ to the 90\u25e6 rotate scanning strategy, the shape of the external and internal contours fabricated by 67\u25e6 rotate scanning strategy are circle approximately. This is caused by that the 67 is a prime number, the first and last track can occur at any positions when the track number is large enough. In view of the overall situation, the step effect cannot occur and the dimension distribution is nearly uniform. But the shape distortion is still existed in the adjacent layers, as shown in Fig. 3. The model of the SLMed circle structure with \u201cskin-core\u201d scanning strategy is shown in Fig. 4. As can be seen in Fig. 4, the track by the \u201cskin\u201d scanning will cover the external contour and internal contour in every layer. Thus, no matter the rotate angle of the \u201ccore\u201d scanning is, the shape of the external and internal contours are standard circle. The step effect and shape distortion will not existed and the dimension distribution is uniform. Consequently, the SLMed circle structure by the \u201cskin-core\u201d scanning strategy has a better geometric accuracy rather than by the 90\u25e6 and 67\u25e6 rotate scanning strategies"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001583_1077546320916628-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001583_1077546320916628-Figure2-1.png",
        "caption": "Figure 2. Pantograph dynamic operation test bench structure diagram.",
        "texts": [
            " Put the EEMD-entropy features of the training samples into PSO\u2013SVM and obtain the PSO\u2013SVM model. 4. Put the EEMD-entropy features of the testing samples into the optimized PSO\u2013SVM model to verify the effectiveness of the diagnosis method. 5. Form high dimensional feature set from EEMDentropy features and use multiple feature ranking criteria to select the feature and enhance the fault diagnosis accuracy. All the vibration data in this article come from a test bench with CED100 pantograph as in Figure 2. The test bench can simulate the pantograph\u2013catenary interaction under the actual running state of the vehicle, with the Z font movement of the contact line and vertical vibration of the pantograph base to achieve a more accurate vibration data. The experimental vibration data are obtained from four pantograph fault types which are healthy, horn crack fault, head bracket crack fault and composite fault. The fault components are shown in Figure 3. The length of head bracket crack is 30 mm and the length of horn crack is 15 mm"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001412_j.mechmachtheory.2019.103730-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001412_j.mechmachtheory.2019.103730-Figure5-1.png",
        "caption": "Fig. 5. A line-symmetric-based 7R mechanism with a tangential intersection in the configuration space.",
        "texts": [
            " Nevertheless, this situation can be identified when using the analysis method described in Section 4 , not only does the method reveal the branching at the solution set of the order higher than the order of contact, it also shows that the time derivative of this order of the joint variable q 7R is different to 0, indicating that the joint is not inactive. 6. Examples In this section two examples of reconfigurable mechanisms with two branches of motion intersecting tangentially are presented. Both examples are obtained using the method discussed in Section 5 . The tangential intersection is identified using the concepts discussed in Section 4 . See [31,33,57\u201359] for more detailed examples of the computation of the solutions to the higher-order kinematic analyses. 6.1. Case 1: a line-symmetric-based 7R mechanism Fig. 5 a shows a 7R mechanism in which joints with axes S 1 , S 2 , S 1 , S 3 , S 4 and S 2 constitute a Bricard linesymmetric 6R linkage [61,62] . As shown in Fig. 5 b, in the configuration q 0 \u2208 V , axes S 1 , . . . , S 4 lie on plane , while axes S 1 and S 2 lie on and are perpendicular to . A seventh R joint is inserted between joints with axes S 1 and S 2 . At q , the axis of the seventh joint, S , is parallel to S and also lies on . 0 7R 1 The screw coordinates with respect to the coordinate system with origin at O shown in Fig. 5 a are the following: S 1 ( q 0 ) := ( 0 , \u2212 1 , 0 ; 0 , 0 , 0 ) , S 2 ( q 0 ) := ( \u22121 , 0 , 0 ; 0 , 0 , 0 ) , S 1 ( q 0 ) := ( 0 , 0 , 1 ; 1 , 0 , 0 ) , S 3 ( q 0 ) := ( 0 , 1 , 0 ; 0 , 0 , 2 ) , S 4 ( q 0 ) := ( 1 , 0 , 0 ; 0 , 0 , \u2212 1 ) , S 2 ( q 0 ) := ( 0 , 0 , 1 ; 0 , \u2212 2 , 0 ) , S 7R ( q 0 ) := ( 0 , 0 , 1 ; 1 2 , \u2212 1 , 0 ) , Define x i := (x 1 i , . . . , x 7 i ) \u2208 R 7 , where x 1 i := d i q 1 / d t i , x 2 i := d i q 2 / d t i , x 3 i := d i q 1 / d t i , x 4 i := d i q 3 / d t i , x 5 i := d i q 4 / d t i , x 6 i := d i q 2 / d t i and x 7 i := d i q 7R / d t i "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002145_978-981-15-5580-0-Figure9.1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002145_978-981-15-5580-0-Figure9.1-1.png",
        "caption": "Fig. 9.1 Design-functional diagramof integrated sensor system for controlling the altitude\u2013velocity parameters of unmanned AP based on the vortex method",
        "texts": [
            " Amplitude of pressure pulsationsPm calculated for different altitude and velocities of incoming air flow to wedge-shaped pyramid is shown in Table 9.1. Bold combination altitude and velocities of flight for manned and unmanned subsonic aircraft plane are not working. The use of the vortex method allowed the authors to develop an original designfunctional diagramof the sensor system for controlling the altitude\u2013speed parameters of an unmanned AP with a single integrated receiver of primary information, shown in Fig. 9.1 [7]. The integrated sensor system of altitude\u2013velocity parameters of the unmanned AP contains two wedge-shaped pyramids 1, the bases of which located orthogonal to each other and counter the incoming air flow. The velocity vector of the incoming air flow V is equal in magnitude and has negative sign to the vector Va of the true airspeed of unmanned AP. Angle of direction of flow is equal in magnitude and sign of measured aerodynamic angle, for example, the incidence angle \u03b1 of unmanned AP. Wedge-shaped pyramids are installed on board of the unmanned AP so that the axis of the wedge-shaped pyramids will be perpendicular to plane of changes of the measured incidence angle \u03b1",
            " The pneumo-electric converters 2 are connected to the recording devices 3 of frequency, which measure the frequencies f 1 and f 2 of vortex formation behind the wedge-shaped bodies. Measured frequency f 1 and f 2 fed to the input of the processing device 4, in which digital signals (codes) are determined and formed according to the developed algorithms according to the value of the true airspeed VB and aerodynamic angle \u03b1. When the constructive implementation of the integrated sensor system of the unmanned AP the wedge-shaped pyramids are installed coaxially above each other, as shown in Fig. 9.1. To ensure stable vortex formation and eliminate the influence of the skewing of the incoming air flow in a plane is perpendicular to the vertical axis of wedge-shaped pyramids on the lower and upper surfaces of pyramids the straightening vanes 5 are installed which made in the form of thin disks. These disks allocate zones with stable vortex formation behind the wedge-shaped pyramids in the incoming air flow and reduce errors caused by the skewing of the flow in the plane perpendicular to the plane of measurement",
            " The researches have shown [9] that the frequencies of vortex formation behind wedge-shaped pyramids, located by the angle 2\u03c60 = 90\u00b0, are determined by the relations: f1 = Sh l Va sin(\u03c60 + \u03b1) = \u221a 2 Sh l Va cos\u03b1 + sin \u03b1 ; f2 = Sh l Va sin(\u03c60 \u2212 \u03b1) = \u221a 2 Sh l Va cos\u03b1 \u2212 sin \u03b1 , (5) where Sh\u2014number Struhal\u2019s of wedge-shaped pyramid, l\u2014size of base. To calculate the true airspeed VB and aerodynamic angle \u03b1 with use of integrated sensor system for controlling the altitude\u2013velocity parameters of unmanned AP, the authors developed original processing algorithms of frequencies of vortex formation of the form [8]: Va = l\u221a 2Sh f1 f2\u221a f 21 + f 22 ; \u03b1 = arctg f2 \u2212 f1 f1 + f2 (6) To increase the functionality and measuring all altitude\u2013velocity parameters of unmanned AP, the authors proposed [10] on the surface of the upper straightening vane 5 (Fig. 9.1) install the receiver 6 for the perception of static pressure PH of incoming air flow. Receiver 6 through a pneumatic duct 7 is connected with pneumoelectric sensor 8 of absolute pressure with the frequency output signal. The output of the pneumo-electric sensor 8 in the form of the frequency fPH is proportional to the static pressure PH of the incoming air flow connected to the input of the processing device 4. The processing device 4 is made in the form of a computer, which according with developed algorithms by the authors calculating and formatting output digital signals of integrated sensor system for controlling altitude\u2013velocity parameters of unmanned AP"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000245_j.jmatprotec.2019.116515-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000245_j.jmatprotec.2019.116515-Figure9-1.png",
        "caption": "Fig. 9. Flange wrinkling model in SSFC spinning: a) simulation result; b) compressive stress ring.",
        "texts": [
            " Measurement results show that in the SSFC spinning the peak position of the wrinkle waveform is similar to that of the conventional spinning, while the waveform valley is basically parallel to the horizontal line, which is due to the limitation of the backup roller. The wrinkle waveforms cannot be extended downward due to the constraint of the backup roller, therefore, the deflection mode in the SSFC spinning is assumed as follows: = +w w m\u03b8 f r[1 cos( )] ( )0 (4) where, m\u03b8cos ( ) describes the deflection mode in the circumferential direction; f r( ) denotes a function of the variable r , which is used to describe the deflection mode in the radial direction. Fig. 9 is the distribution of circumferential compressive stress in the blank material when the SSFC spinning angle is 35\u00b0, where r1 and r2 are the inner and outer radii of the deformed flange, respectively; the point S is the node with maximum circumferential compressive stress; the Table 2 Mechanical properties of 2024-O aluminum alloy. Material Young\u2019s modulus (GPa) Poisson\u2019s ratio Yield strength (MPa) Strength coefficient (MPa) Hardening exponent 2024-O 71.3 0.33 70.06 308.62 0.23 points M and N are intersections between the rays passing through the point S and edges of the stress ring, respectively",
            " Under the condition of =\u0394U \u0394T , the critical circumferential compressive stress \u03c3\u03b8cri can be calculated by means of numerical method. During spinning process, if the absolute value of the maximum circumferential compressive stress \u03c3\u03b8max in the deformed flange is greater than that of the critical circumferential compressive stress \u03c3\u03b8cri at the corresponding time, the flange wrinkling will initiate. In order to calculate the stress \u03c3\u03b8cri accurately, it is necessary to combine finite element simulation to extract the stress \u03c3r and the radii r1 and r2 of the deformed flange. As shown in Fig. 9, after the coordinates, circumferential compressive stress and radial tensile stress of the three points of M, S and N were extracted respectively from the spinning simulation, the radii r1 and r2 can be calculated according to the coordinates of M and N. In addition, it should be pointed out that the thickness of the stress ring can be considered as the initial value of the blank material. Fig. 10 is a wrinkling prediction result for the 2024-O hemispherical part SSFC spinning based on the developed prediction model, where SSFC-Max represents the maximum circumferential compressive stress \u03c3\u03b8max obtained from spinning simulation; SSFC-Cri represents the critical circumferential compressive stress \u03c3\u03b8cri calculated by the prediction model"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001001_sta.2019.8717197-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001001_sta.2019.8717197-Figure7-1.png",
        "caption": "Fig. 7. Bearing fault detection test bench.",
        "texts": [
            " Initially, the vibration signals are collected at an fs of 48 KHz, then resampled to an fs of 1860 Hz in order to facilitate the wavelet decomposition. Actually, the signals are resampled via the resampling MATLAB block, which contains a digital anti-aliasing filter. The Op-SWPT calculates the coefficient containing the bearing fault related frequency, then, the Root Mean Square (RMS) of the extracted coefficient is computed in order to construct training and testing datasets. Finally, a Decision Tree classifier classifies the data. The experimental bench of the CWRU is presented in Fig.7. It consists of an IM of 2 hp with a torque encoder and transducer. The resistant load is applied to the IM by a dynamometer controlled by an electronic system. The tests explored in this work include 6205 bearings. Table. I lists the bearing characteristics and the fault frequencies corresponding to the tested bearings. Faults in the OR of the drive end bearing with holes diameters varying from 0.007 to 0.028 inches are selected. Accordingly, four motor conditions are studied: a healthy motor (HLT), a motor with an OR fault with a 0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003900_jsvi.1997.1323-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003900_jsvi.1997.1323-Figure2-1.png",
        "caption": "Figure 2. Model of the flexible sleeve.",
        "texts": [
            " The chamber pressure can also be changed dynamically by a servo valve. So the active journal bearing can deliver dynamic control forces to the rotor via the oil film to control the forced vibration by either an open-loop means or feedback approaches. To obtain the stiffness and mass matrices of the flexible sleeve, the finite element method (FEM) was employed. The flexible sleeve was considered as a curved cantilever beam. Each node has three degrees of freedom (d.o.f.) s, n, y, which correspond to nodal forces S, N and M as shown in Figure 2. By assembling the individual elements along the global co-ordinates xg , yg and vg , the initial model of the flexible sleeve was obtained. In order to reduce the number of d.o.f., the Guyan reduction technique [14] was used to condense the mass and stiffness matrices. The condensation was performed in three steps. First, the original matrices were condensed along co-ordinates xg and yg (the angular co-ordinates were eliminated). Then, the matrices were transferred to a system of co-ordinates r, t"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002042_j.addma.2020.101730-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002042_j.addma.2020.101730-Figure9-1.png",
        "caption": "Fig. 9. Scheme of the crack front shape evolution during crack propagation in the first two iterations (a) and a detailed view of the finite element model showing the very fine mesh surrounding the crack front (b).",
        "texts": [
            " The next step in the modelling was the determination of the local stress intensity factor KI as a function of crack growth (\u0394a), which is necessary for the lifetime estimation. The crack propagation in the model was implemented in the form of iterations. According to Refs. [9, F. Arbeiter et al. Additive Manufacturing 36 (2020) 101730 58], the size of printing defects in well printed FFF parts are usually in the range of 50 and 100 \u00b5m. Similar defect sizes were found on fracture surfaces in the present work (see magnification in Fig. 7). Therefore, the crack length for the first iteration was set to 50 \u00b5m with a semi-circular crack shape (Fig. 9a). Here, KI was calculated along the crack front and evaluated for two points. The first point (A1) was the deepest point of the crack front (point on the symmetry plane). The second point (X1) lies on the crack front with a distance A1\u2013X1 of approximately 70% of the whole crack front in order to be far enough from the free surface, where the stress intensity factor concept fails due to an additional singular field in the corner [48]. As the first increment of the semi-axis, a was chosen as \u0394a1 = 0",
            " a2 = a1 +\u2206ai (6) x2 = x1 +\u2206xi (7) Subsequently, by these two points A2 and X2, the ellipse representing the crack front is fitted and the dimension of the semi-axis b is obtained. This procedure was repeated with a constant increment of the semi-axis a of ai = 0.1 mm. All configurations, for which the stress intensity factor was calculated, are shown in Table 3. When the size of the simulated crack becomes similar to the width of the half model, meshing of the crack front becomes increasingly difficult (Fig. 9b) and requires significant higher calculation times. Furthermore, it should be mentioned that also the real crack front is flattened during propagation, as shown in Fig. 7. Therefore, the crack propagation was not simulated with this 3D model throughout the whole width of the component. For crack lengths longer than 1 mm, a 2D model was used. In this way, calculation time of approximately one order of magnitude was saved. The 2D model was created similarly to the 3D model with some minor modifications (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000275_tie.2019.2959496-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000275_tie.2019.2959496-Figure2-1.png",
        "caption": "Fig. 2. Cross sectional view of the solid rotor (dimensions in mm)",
        "texts": [
            " Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. determined using (2): v = iR+ d\u03c8 dt (1) \u03c8(t) = \u222b t t0 (v \u2212 iR)d\u03c4 + \u03c8(t0) (2) where v and i are phase voltage and current, respectively. A representative sectional view of a 6/4 SRM is shown in Fig. 1. The geometric details of the stator of the machine under test are given in Table I. The cross-sectional view of the solid-rotor is given in Fig. 2. Instead of solid EN-8 material, if the rotor were made of laminations of the same material, the machine flux-linkage characteristics would appear as shown in Fig. 3. These static characteristics are obtained through 2D finite element analysis. The measured static B-H curves of the stator and rotor materials [31] are used in this FE analysis. Determination of the flux linkage characteristics of the solidrotor SRM is discussed below. The solid-rotor SRM under test (photograph shown in Fig. 4) is fed with voltage pulses using an asymmetric H-bridge converter (Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003126_s42835-021-00852-z-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003126_s42835-021-00852-z-Figure6-1.png",
        "caption": "Fig. 6 The rectangular conductor is in a uniformly distributed transverse sinusoidal magnetic field",
        "texts": [
            " A heat generation per unit volume of the conductor needs to be calculated due to the flow of current through the conductor for the thermal analysis. The stator losses mainly consist of wingding loss, iron loss, and additional loss while winding loss can be divided into eddy current loss and copper loss. It is assumed that the individual strands ensure that the current distribution in the conductors is uniform and therefore does not contribute to additional losses. A rectangular conductor in a magnetic field is shown in Fig.\u00a06. Especially, the magnetic field B\u0307 0 is assumed to be a uniformly distributed sinusoidal magnetic field and parallel to the x-axis. Around the neutral plane ( y = 0 ), electric field intensity E\u0307Z and eddy current will be induced in the rectangular conductor. It is known that [22] 1 3 where, is the frequency of the magnetic field B\u0307. Then, Therefore, the amplitude of the electric field intensity E\u0307Z distance from the neutral plane y is The amplitude of the density J z is as follows where, is the electrical conductivity of the conductor",
            " Each turn cooper transposed conductor is built as an entity with an anisotropic equivalent Young\u2019s modulus by the method of representative volume elements [24]. The outer surface of the back iron is fixed and the two parallel symmetry faces could translate in the X and Y directions, while fix in the Z direction. The defined direction is similar as shown in Fig.\u00a03. The twelve conductors with two columns in the slot region are numbered 1 to 6 correspond with the EM modal. By loading the EM forces from the EM analysis (results summarized in Table\u00a02) and the temperature distribution from the CFD analysis (as shown in Fig.\u00a06). The Von-Mises stresses contour in the slot region of the stator is shown in Fig.\u00a08. The contour indicates that there is a stress concentration. The maximum stress is about 290\u00a0MPa and it is located in the bottom corners of the GFRP teeth because of the different thermal expansion coefficients between the copper conductor and the GFRP teeth. Especially, the tension strength of the GFRP is about 452\u00a0MPa [25]. Therefore, the safety factor is about 1.6 (452/290 = 1.6), then the GFRP teeth meet the strength requirement of the HTS generator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001326_00207721.2019.1692094-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001326_00207721.2019.1692094-Figure1-1.png",
        "caption": "Figure 1. Geometric meaning of the unicycle control laws in (7) and (8).",
        "texts": [
            " However, in the leader\u2013follower manoeuvre control, the leaders\u2019 velocities are generally time varying, then the integral part is necessary for eliminating the steady-state tracking error. Comparing with other intelligent control methods, the proposed control laws are distributed and easy to implement. Remark 3.1: Since each agent only requires the relative position measurements {pi \u2212 pj}j\u2208Ni of its neighbours, the proposed control law is distributed. Second, the geometric interpretation of the proposed control laws in (7) and (8) is shown in Figure 1. The linear and angular velocities of the agents are equal to the magnitudes of the orthogonal projection of \u2211 j\u2208Ni Pg\u2217 ij (pi \u2212 pj) onto \u03c7\u03b8i and \u03c7\u22a5 \u03b8i , respectively. The similar geometric control design approach is also used in Zhao and Zelazo (2017) and Zhao et al. (2018). Different from the control law in Lin, Francis, and Maggiore (2005), the angular control law is orientational in this paper. Moreover, the heading vector \u03c7\u03b8i is very flexible. In this way, the bearing-based formation control problem can be solved and, in the meantime, other tasks such as obstacle avoidance can also be achieved"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002065_j.matpr.2020.10.324-Figure2-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002065_j.matpr.2020.10.324-Figure2-1.png",
        "caption": "Fig. 2. FEA of Spring 1.",
        "texts": [],
        "surrounding_texts": [
            "In this paper, a helical coil spring made of IS 4454 used in the automotive application has been optimised using multi-objective optimisation procedure. Bi-objectives namely, minimizing the volume and maximizing the strain energy were considered in this study along with realistic constrains. Non-dominated pareto solutions were generated using NSGA II. A total of 18 non-dominated solutions were generated by NSGA-II. To validate the solutions generated from NSGA II, FEA has been carried out by selecting extreme solutions of the Pareto front and middle one. Strain energy and volume of the springs were computed using the FEA procedure, and the results were compared with the pareto solutions generated by NSGA II. It is observed that there is a close correlation exists between the solution obtained by NSGA II and numerical procedure. A maximum relative percentage error of 4.88% was observed with respected to strain energy. The results show the effectiveness of NSGA II. The designer can choose any one of the non-dominated solution obtained by NSGA II for the automotive application based on their priority. Table 8 FEA results comparison. Spring Model NSGA II Results FEA results Relative % error(Strain Energy) Volume (mm3) Strain energy (Joules) Volume (mm3) Strain energy (Joules) 1 228742.70 327046.31 228742.70 335790.00 2.67 2 242262.00 371982.43 242262.00 390150.00 4.88 3 261081.80 405378.12 261081.80 412550.00 1.77 Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper. References [1] S. Kirkpatrick, Optimization by simulated annealing: Quantitative studies, J. Stat. Phys. 34 (1984) 975\u2013986. [2] F. Glover, Tabu Search and Adaptive Memory Programming Advances, Applications and Challenges, in: Interfaces in Computer Science and Operations Research, Springer, Boston, MA, 1997, pp. 1\u201375. [3] D.E. Goldberg, J.H. Holland, Genetic algorithms, and machine learning, Machine Learn. 3 (1988) 95\u201399. [4] M. Dorigo, T. St\u00fctzle, Ant colony optimization: overview and recent advances. In Handbook of metaheuristics, Springer, 2019, pp.311-351. [5] R. Storn, K. Price, Differential evolution\u2013a simple and efficient heuristic for global optimization over continuous spaces, J. Global Optim. 11 (1997) 341\u2013 359. [6] R. Eberhart, J. Kennedy, A new optimizer using particle swarm theory, In MHS\u201995. Proceedings of the Sixth International Symposium on Micro Machine and Human Science, IEEE, 1995, pp.39-43. [7] E. Sandgren, Nonlinear integer and discrete programming in mechanical design optimization, J. Mech. Des. 112 (1990) 223\u2013229. [8] K. Deb, Optimal design of a welded beam via genetic algorithms, AIAA J. 29 (1991) 2013\u20132015. [9] K. Deb, M. Goyal, Optimizing engineering designs using a combined genetic search, ICGA (1997) 521\u2013528. [10] Y.J. Cao, Q.H. Wu, A mixed variable evolutionary programming for optimisation of mechanical design, Eng. Intell. Syst. Electric. Eng. Commun. 7 (1999) 77\u201382. [11] S. He, E. Prempain, Q.H. Wu, An improved particle swarm optimizer for mechanical design optimization problems, Eng. Optim. 36 (2004) 585\u2013605. [12] R. Roy, S. Hinduja, R. Teti, Recent advances in engineering design optimisation: Challenges and future trends, CIRP Ann. 57 (2008) 697\u2013715. [13] K. Rameshkumar, R.K. Suresh, K.M. Mohanasundaram, Discrete particle swarm optimization (DPSO) algorithm for permutation flowshop scheduling to minimize makespan, Lecture Notes Comput. Sci. 3612 (2005) 572\u2013581. [14] K. Rameshkumar, C. Rajendran, K.M. Mohanasundaram, A novel particle swarm optimisation algorithm for continuous function optimisation, Int. J. Operat. Res. 13 (2012) 1\u201321. [15] K. Rameshkumar, C. Rajendran, K.M. Mohanasundaram, Discrete particle swarm optimisation algorithms for minimising the completion-time variance of jobs in flowshops, Int. J. Ind. Syst. Eng. 7 (2011) 317\u2013340. [16] K. Rameshkumar, C. Rajendran, A novel discrete PSO algorithm for solving job shop scheduling problem to minimize makespan, IOP Conf. Ser. Mater. Sci. Eng., 310 (1), 2018, art. no. 012143. [17] R.V. Rao, V.J. Savsani, Mechanical Design Optimization Using Advanced Optimization Techniques, Springer Science & Business Media, 2012. [18] M. Taktak, K. Omheni, A. Aloui, F. Dammak, M. Haddar, Dynamic optimization design of a cylindrical helical spring, Appl. Acoust. 77 (2014) 178\u2013183. [19] T.P. Bagchi, Multiobjective scheduling by genetic algorithms, Springer Science & Business Media, 1999. [20] K. Deb, S. Agrawal, A. Pratap, T. Meyarivan, A fast elitist non-dominated sorting genetic algorithm for multiobjective optimization: NSGA-II. In International conference on parallel problem solving from nature Springer, Berlin, Heidelberg, 2000, pp.849-858. [21] K. Deb, Multi Objective Optimization Using Evolutionary Algorithms, John Wiley and Sons, 2001. [22] C. Coello, S. De Computaci\u00f3n, C. Zacatenco, Twenty years of evolutionary multi-objective optimization: A historical view of the field, IEEE Comput. Intell. Mag. 1 (2006) 28\u201336. [23] M. Reyes-Sierra, C. Coello, Multi-objective particle swarm optimizers: A survey of the state-of-the-art, Int. J. Comput. Intell. Res. 2 (2006) 287\u2013308. [24] K.E. Parsopoulos, M.N. Vrahatis, Multi-objective particles swarm optimization approaches. In Multi-objective optimization in computational intelligence: Theory and practice, IGI global, 2008, pp.20-42. [25] N. Gunantara, A review of multi-objective optimization: Methods and its applications, Cogent Eng. 5 (2018) 1\u201316. [26] P. Sabarinath, M.R. Thansekhar, R. Saravanan, Multiobjective optimization method based on adaptive parameter harmony search algorithm, J. Appl. Math., (2015). [27] M. Taktak, F. Dammak, S. Abid, M.A. Haddar, Finite element for dynamic analysis of a cylindrical isotropic helical spring, J. Mech. Mater. Struct. 3 (2008) 641\u2013658. [28] T. Habuchi, H. Tsutsui, S. Tsuji-Iio, R. Shimada, Optimization of the stress distribution of helical coils with a cable-in-conduit configuration, IEEE Trans. Appl. Supercond. 21 (2011) 3494\u20133500. [29] T.S. Sarkate, A finite element approach for analysis of a helical coil compression spring using CAE tools, Appl. Mech. Mater. 330 (2013) 703\u2013707. [30] B.L. Choi, B.H. Choi, Numerical method for optimizing design variables of carbon-fiber-reinforced epoxy composite coil springs, Compos. B Eng. 82 (2015) 42\u201349. [31] H.B. Pawar, D.D. Desale, Optimization of three-wheeler front suspension coil spring, Procedia Manuf. (2018) 428\u2013433. [32] P. Dhanapal, K.R. Mathevanan, M.R. Krishnan, R. Thennarasan, Design and analysis of composite helical spring for two-wheeler shock absorber, Int. J. Pure Appl. Math. 118 (2018) 549\u2013555."
        ]
    },
    {
        "image_filename": "designv11_14_0000380_978-981-13-6647-5_10-Figure10.36-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000380_978-981-13-6647-5_10-Figure10.36-1.png",
        "caption": "Fig. 10.36 The structure of the Hao fine-breaking machine [2]: 1\u2014beater roll, 2\u2014blade bottom, 3\u2014baffle, 4\u2014top cover, 5, 6\u2014circulation groove, and 7\u2014butted-knife installation",
        "texts": [
            " The chopping result by dick mill is next to the result of Hao fine-breaking machine. Increase the chopping time by disk mill form 9 times (curve 2) to 12 times (curve 4) significantly improve the even distribution of fiber length and reduce the content of long fiber. The results of the test are shown in Table 10.42. Commonly used fine-breaking equipment is Hao fine-breaking machine. Continuous fine-breaking equipment includes a cone-shaped chopper, cylindrical chopper, and disk mill. (1) Hao fine-breaking machine The structure of the Hao fine-breaking machine is shown in Fig. 10.36. Hao fine-breaking machine is characterized by easy control of the process conditions, which is suitable for a variety of different requirements of the fine pulp. However, it has some drawbacks, such as low production capacity (volume of 400\u2013500 kg), difficulty in control the quality of each batch, and high labor intensity. In the process of fine-breaking, turn the hand wheel of the butted-knife installation (7) to adjust the height of the beater roll (1), and the knife distance is based on the loaded current of the motor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000683_tmag.2016.2601886-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000683_tmag.2016.2601886-Figure3-1.png",
        "caption": "Fig. 3. Topologies of SS-PMM and two proposed PS-PMMs with 12/10 stator/rotor pole number combination.",
        "texts": [
            " Secondly, when the ratio of stator pole number to the greatest common divisor of stator and rotor pole number is equal to even integers, bipolar phase flux-linkage and symmetrical phase back-EMF waveforms can be obtained, although the coil flux-linkage and coil back-EMF are unipolar and asymmetric due to the existence of even harmonics [21], [28]. Thirdly, the reluctance torque is negligible [21], [28]. Moreover, due to the same operational principle, the coil EMF phrasor method which used to determine the winding configuration of SS-PMMs can be extended to PS-PMMs, where the electrical degree between two adjacent coil-EMF phrasors can be calculated from the mechanical degree and rotor pole number [28]. By way of example, for the 12/10 stator/rotor pole (12S/10R) SS-PMM, PS-PMM-I and PSPMM-II as shown in Fig. 3, the corresponding coil-EMF phrasors are shown in Fig. 2. Accounting for the alternate magnetization directions in adjacent stator poles, coil ni and coil ni' are referred as the coils with additional opposite polarities, such as coils 1 (A1) and 2 (B4) as shown in Fig. 2(b) and Fig. 3. PS-PMMS In this section, the electromagnetic performance of the proposed PS-PMM-I and PS-PMM-II will be analyzed and compared with the original SS-PMM under the same 12/10 stator/rotor pole combination and the same machine size as well as the same rated copper loss. Then, the influence of stator and rotor pole number combinations on electromagnetic performance of PS-PMM-Is and PS-PMM-IIs will also be analyzed. All machines are globally optimized with the objective of maximum average torque under the same rated 30W copper loss by genetic algorithm (All of the geometric parameters are considered in the global optimization)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000682_ufne.2016.01.037677-Figure5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000682_ufne.2016.01.037677-Figure5-1.png",
        "caption": "Figure 5. Total potential U of elastic and van der Waals forces, U Uelastic UVdW, calculated at a varied distance z0 (a) and coefficient of elasticity K (b). Arrows indicate the direction of z0 and K growth.",
        "texts": [
            " In the numerical model [14], the fractal surface can be given in the form of an x-coordinate-dependent data block; in accordance with the standard definition w x 1 2p qmax qmin B q cos qx z ; 5 where Fourier series coefficients have the scale-invariant form,B q c0q a, and z x is the d-correlated randomphase: z x z x 0 d x\u00ff x 0 : 6 Certainly, the real surface w x extending from zero distances to infinity can never be truly fractal; it is so-called quasifractal including a certain limited spectrum of wave vectors qmin < q < qmax in formula (5). Its maximum and minimum amplitudes (i.e., roughness or, in the mathematical language, standard deviation) are limited, too:D\u00ff w x \u00ff w x 2E1=2 4A ; 7 where parameter A describes the characteristic physical roughness of the surface. In a certain region of parameters, the total potential containing adhesive and elastic components has two valleys (wells) of comparable depths. This total potential is presented in Fig. 5 in the form of families of curves calculated at different distances and elastic constants. Evidently, there are in both cases parameter regions, where the potential passes through two wells. In accordance with general principles of physical kinetics, fluctuating parameters can be expected to give rise to two alternative states of the system, with comparable energy valleys provoking dynamic jumps of the fiber ends between two (attached and detached) states. Such peculiar `exchange of excitations' averaged over time assures attraction to the surface [14]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001955_j.mechmachtheory.2020.104090-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001955_j.mechmachtheory.2020.104090-Figure9-1.png",
        "caption": "Fig. 9. Figure illustrating the surfaces \u03c8 i (c 1 , c 2 , c 3 ) = 0 , i = 1 , 2 , 3 along with d 234 = 0 corresponding to Fig. 8 .",
        "texts": [
            " 080 c 1 c 2 + 0 . 745 c 1 c 3 + 0 . 496 c 1 \u2212 0 . 001 c 2 2 \u2212 0 . 106 c 2 c 3 \u2212 0 . 143 c 2 + 0 . 137 c 2 3 + 0 . 306 c 3 + 0 . 158 = 0 . (41) This example clearly demonstrates the utility of the analysis of the extra factors. As the MP has a special geometry in this case, certain combinations of planes (i.e., { P 21 , P 31 , P 41 }, in this case) fail to intersect at a point, but only for a set of discrete points in SO (3) , which, in turn, lie on the surface in SO (3) given by Eq. (41) (see Fig. 8 ). In Fig. 9 , these points (i.e., only of pair of them, for the sake of visual clarity) are shown to lie at the intersections of the surfaces \u03c8 i = 0 , i = 1 , 2 , 3 , with the surface d 234 = 0 . These points are extremely special, as they represent the solutions of an overconstrained system containing these four polynomials in { c 1 , c 2 , c 3 }. Without the knowledge of these points, the standard formulation of forward kinematics would suffer an unexplained failure at these combinations of planes and orientations"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001331_j.matpr.2019.11.030-Figure3-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001331_j.matpr.2019.11.030-Figure3-1.png",
        "caption": "Fig. 3. Un-deformed and deformed three-point bending test configuration.",
        "texts": [
            " The test process involves placing the test specimen in the testing machine and applying tension to it until it fractures. During the application of tension, the elongation of the gauge section is recorded against the applied force. Please cite this article as: M. Vineeth Choudary, A. Nagaraja, K. Om Charan Sai e fibre, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2019.11.03 The aim of this test is to determine the value of Young\u2019s modulus of the specimen using three-point bend test as shown in Fig. 3. The specimens are in ASTM D790 dimension. The modulus of elasticity is calculated by steepest initial straight-line portion of the load-deflection curve. Glass fibre reinforced plastics (GFRP) are used on a large scale in aeronautical industry due to their advantages. In the same time, GFRP can present degradations during their use, as delimitation due to some impact even with low energies, accompanied or not by fibres breaking, local overheating, water adsorption, and the last two causes leading to the deterioration of the matrix"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000597_s11665-016-2214-1-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000597_s11665-016-2214-1-Figure6-1.png",
        "caption": "Fig. 6 Shape of the double ellipsoid heating source employed in the FEM simulations",
        "texts": [
            " Concerning the shape of the heating source, the most common source employed to simulate keyhole LBW is the conical Gaussian distribution (Ref 19). However, it was experimentally observed that the heating source model that better fitted to the real shape of the HPDL radiation (provoking welding by conduction regime, as explained before) was the \u2018\u2018Goldak double ellipsoid model,\u2019\u2019 available within the SYSWELD code. The double ellipsoid model is based on the work of Goldak (Ref 27). The shape of this heat source is shown in Fig. 6. This heating source has been already employed to successfully simulate bead-on-plate TIG welds (Ref 28), but it has not been previously used (as far as the authors are concerned) to simulate conduction LBW. Therefore, the present research study is the first work in which this heating source model is applied to simulate LBW of Ti6Al4V under conduction regime. To obtain a reasonable agreement, the size and morphology of the heating source and the laser efficiency were estimated by reiterative refinement steps"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002352_j.mechmachtheory.2020.103945-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002352_j.mechmachtheory.2020.103945-Figure8-1.png",
        "caption": "Fig. 8. Two \u201c{6R} [2R] (5R)\u201d deployable units and two \u201c{6R} [1R] (6R)\u201d deployable units",
        "texts": [
            " 7 (a) and (b). Similarly, seventeen available \u201c{6R} [3R] (4R)\u201ddeployable units can be obtained, listed in Table 3 , in which two typical units are shown in Fig. 7 (c) and (d). Based on the combination of three kinds of \u201c{6R}\u201d modules, one kind of \u201c[2R]\u201d module and three kinds of \u201c(5R)\u201d modules, there are nine \u201c{6R} [3R] (5R)\u201d units in total. After excluding those units with structural defects, six available\u201c{6R} [2R] (5R)\u201d units can be obtained, listed in Table 4 , in which two typical units are shown in Fig. 8 (a) and (b). Similarly, there are also six available \u201c{6R} [1R] (6R)\u201ddeployable units, listed in Table 5 , in which two typical units are shown in Fig. 8 (c) and (d). Please cite this article as: W.-a. Cao, Z. Jing and H. Ding, A general method for kinematics analysis of two-layer and twoloop deployable linkages with coupling chains, Mechanism and Machine Theory, https://doi.org/10.1016/j.mechmachtheory. 2020.103945 6 W.-a. Cao, Z. Jing and H. Ding / Mechanism and Machine Theory xxx (xxxx) xxx An umbrella-shaped deployable mechanism can be constructed by connecting a serial of identical deployable units through some assembly methods[34]. For example, by sharing a FB and connecting every two units through an R pair, six \u201c \u02dc R l1 \u22a5 \u2194 R l2 \u02dc R c1 / \u0303 R c2 \u02dc R u 1 \u22a5 \u2194 R u 2 \u22a5 \u0303 R u 3 \u201d deployable units can construct an umbrella-shaped deployable mechanism, whose folded configuration and deployed configuration are shown in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002825_lcsys.2021.3050093-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002825_lcsys.2021.3050093-Figure1-1.png",
        "caption": "Fig. 1. Schematic representation of three pendulums with length Li and mass mi , connected by horizontal springs Si attached to the suspensions pints at distance li , for i = 1, 2, 3.",
        "texts": [
            " Proposition 3: Consider the multi-agent system (1). Assume that there is a robust inverse optimal local controller (9) for each agent, with V(xi,t) given by (14) such that the inequality (17) holds for the condition (18). Then, the dynamics of each agent i \u2208 V with the local control signal defined in (9), is ISS with respect to the perturbation f\u0304i. Proof: This follows as a consequence of Theorem 2, by considering the specific quadratic V(xi,t). Consider the coupled pendula system presented in Figure 1, composed by three pendulums (agents) with length Li and mass mi, for i = 1, 2, 3. The pendulums are coupled by the horizontal springs, Si, which are attached at a distance li from the suspensions points, for i = 1, 2, 3. In the figure, \u03b8i corresponds to the angle between pendulum i and the vertical axis. Consider the tracking problem in which the state of each pendulum, xi,t, is given by the error between the angle \u03b8i and some desired reference ri and its time derivative, that is, xi,t = [ \u03b8\u0303i\u02d9\u0303 \u03b8i ] = [ \u03b8i \u2212 ri \u03b8\u0307i \u2212 r\u0307i ] , x\u0307i,t = [ \u02d9\u0303 \u03b8i \u03b8\u0308i \u2212 r\u0308i ] "
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003833_jsvi.1997.1389-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003833_jsvi.1997.1389-Figure1-1.png",
        "caption": "Figure 1. (a) Schematic of impact/friction oscillator, (b) piecewise linear friction model, (c) trilinear impact model.",
        "texts": [
            " [52] discuss both the effects of friction and impact in experimental systems. The present work includes frictional damping in an extensive study of a two-sided impact oscillator, both in simulation and experiment. It is demonstrated that a relatively simple experiment exhibits the spectrum of non-linear features observed in, and predicted by the mathematical model. Consider a single-degree-of-freedom oscillator consisting of a rotational inertia, harmonically forced through a base spring, and subject to impact at the critical angles fL and fR , as shown in Figure 1(a). Friction is assumed to be present at the pivot. For small f, a summation of moments about the pivot provides the equation of motion: If + cL2 s f + kL2 s f+Lf F(f )+Li G(f)= cLs y\u0307+ kLs y, (1) where F(f ) and G(f) represent non-linear forces due to friction and impact respectively and are considered in more detail later, together with an experimental realization of such a system. The base forcing is harmonic at frequency v and amplitude Y0 : y(t)=Y0 sin (vt). The modelling of friction is an extremely difficult task, influenced as it is by a myriad of different factors, and has been the subject of comprehensive study [25, 30, 37, 40]",
            " The interactive effects of friction forces and system dynamics may be intricate including the possibility of \u2018\u2018frictional memory\u2019\u2019 [25]. For the purposes of the current study (where the global dynamics effects of friction are of primary concern) a relatively simple friction law is adopted. A piece-wise linear Coulomb friction relationship can be written as FA (f )=F sgn (f ), for f $ 0, \u2212aF EFA (f )E aF , for f =0. (2) where the parameter a accounts for a static friction level that differs from kinetic friction. A schematic picture of this model is shown in Figure 1(b). The modelling of impact is somewhat more straightforward than friction although there are again a number of choices to be made. A simple coefficient of restitution has been shown to work well in related studies [16]. However, since the current study involves a rubber ball contacting a flat metal surface there is a finite contact time and hence a change of stiffness model is used here where the equation of motion switches (at impact) to one with a much higher (but still linear) stiffness, i.e., a trilinear stiffness [12, 53, 54]. This characteristic is described mathematically as G(f)=8k Li (f\u2212fR ), 0, k Li (f\u2212fL ), fqfR , fL EfEfR , fQfL , 9 (3) and is also shown schematically in Figure 1(c). Here, the contact stiffness is k at the right and left boundaries fL and fR . It is convenient to reduce the number of parameters in the equation of motion. Time is normalized with respect to the primary natural frequency of the system, vn . A non-dimensional co-ordinate, u(t), and several non-dimensional parameters are defined: vn =zkL2 s /I , t=vn t, u(t)=f(t)Ls /Y0, g=v/vn , h= cLs /2Ivn , v\u0304n =zk L2 i /I , f=F Lf /Ls Y0 k, u c (t)=f c (t)Ls /vn Y0, r= v\u0304n /vn =zk L2 i /kL2 s , sR,L =fR,L Ls /Y0"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003679_s0167-8922(08)70490-5-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003679_s0167-8922(08)70490-5-Figure15-1.png",
        "caption": "Fig. 15 Shell surface temperature. Gasoline engine.",
        "texts": [
            "- 8 - Maximum oil film temperature _ _ _ _ _ Minimum oil film temperature I 1 0 150 , L 100 0 180 360 540 720 Crank angle (deg) Fig. 14 Maximum and minimum oil film temperatures. Gasoline engine. Temperatures. The maximum oil film temperature rise (to 141 deg-C) with respect to the oil inlet temperature of 90 deg-C is 51 deg-C, Fig. 14, which is - as expected - substantially higher than for the Diesel engine bearing. Again the minimum oil film temperature is equal to the oil temperature in the oil groove (109.5 deg-C on the average). The oil film temperature (Fig. 13) and shell surface temperature (Fig. 15) distributions The mathematical model of a thermoelastohydrodynamic calculation method for crank train bearings has been described. The method has been used for the analysis of main bearings of both Diesel and gasoline engines. The following conclusions can be drawn from the results: 0 The consideration of the pressure and temperature dependence of the oil viscosity and density (TEHD) may lead to a temporal (an consequently spatial) shift of the absolute minimum of the oil film thickness compared with EHD results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003484_9781119711230-Figure2.5-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003484_9781119711230-Figure2.5-1.png",
        "caption": "Figure 2.5 Workings of an industrial robot.",
        "texts": [
            " \u2022 Industrial Robot The admin of the robotic arm is controlled via a control- ler embedded with a wireless chip so as to control the arm remotely. The central system enables the admin to manage the settings and set up various modes at which the arm needs to run. The system is connected with the robotic arm that collects data from the system via some protocols like file transfer protocol (FTP) that gives instructions on what task is to be done and how. A representation of the process is shown in Figure 2.5. \u2022 Healthcare Robot The various steps to perform robotic surgery are shown in the flow diagram in Figure 2.6 below. The benefits of robotic surgery is that it can precisely perform the surgery without any extra cuts and cannot make mistakes like human doctors can, as continuous monitoring is done during surgery to check if everything is going well. \u2022 Agricultural Robot Agricultural robots can be very helpful to farmers as they really work very hard to cultivate the crops and don\u2019t think about seasonal weather\u2014from the harsh sun of summers to the chilly winters"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000038_j.apsusc.2019.143594-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000038_j.apsusc.2019.143594-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of selective laser melting (SLM). Based on computer-aided design, planar patterns are drawn using a high-power laser to induce local thermalization. The planar structure is deposited in layers and a 3 dimensional workpiece is manufactured by repeating the planar drawing whilst vertically moving the workpiece-lathe and feeding in the powder.",
        "texts": [
            " As of now, that is why a three-dimensional (3D) printing technologies are considered promising tools for driving future industrialization, so called industry 4.0 [1,2]. Additive manufacturing, as opposed to subtractive methodologies, can be used to develop novel methodologies for synthesizing and fabricating complicated structures from even metallic materials [3\u20136]. Radiative energy sources are used for local deformation (i.e., thermalization) of the target materials during sintering and melting. Selective laser melting (SLM) can be used to fabricate parts with complex geometries directly from powder feedstocks [7,8]. Fig. 1 shows a schematic of the SLM process with a target powder material. SLM is advantageous for manufacturing sophisticated and large-scale parts. Because components are fabricated layer by layer in a bottom-up process, SLM can be used to fabricate free-standing pieces and does not rely on specific molds [9]. A very dense workpiece can be obtained by completely melting the metallic particles then immediately solidifying them, and no post-processing steps are required. Local heating and cooling can potentially be used to make glassy metal as the high cooling rate required for vitrification is obtainable during the SLM process"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001287_8756087919887216-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001287_8756087919887216-Figure1-1.png",
        "caption": "Figure 1. Physical model for case under study.",
        "texts": [
            "16\u201318 A critical literature survey reveals that no contribution is available for mathe- matical formulation of the roll coating process with constitutive equations of Rabinowitsch fluid flow. The aim of this article is to develop the flow pattern of the forward roll coating of a Rabinowitsch fluid along with investigating the effects of physical properties of such a material during the coating process. The paper setup is as follows. We will first present the governing equations followed by problem formulation, numerical solutions, and discussion of the results. Let us consider a laminar, steady flow of an incompressible Rabinowitsch fluid to deposit liquid on a moving substrate. Figure 1 illustrates the region under study. There are two counter rotating rolls with an angular velocityU \u00bc Rx, where R is the roll radius and x is their angular velocity. The plane and the roll are moving linearly with the same velocity with H0 as the separation at the nip. A Rabinowitsch fluid is governed by following equations r: V \u00bc 0 (1) q D V Dt \u00bc r p \u00fer: s (2) where In this model, the non-linear relationship between shear stress and rate of defor- mation for one-dimensional flow can be described as follows sxy \u00fe K\u00f0 sxy\u00de3 \u00bc l @ u @ y (3) In above equation, l is the zero-shear viscosity and K is the non-linear parameter which represents the non-Newtonian fluid behavior"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001708_1350650120929896-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001708_1350650120929896-Figure1-1.png",
        "caption": "Figure 1. Meshed model of axisymmetric asperity in contact with a rigid flat plane.",
        "texts": [
            " The Bilinear Isotropic Strain Hardening Option (BISO) in the ANSYS program is opted to account the elastic-plastic behavior of real materials. The rate-independent plasticity algorithm incorporates the von Mises criterion, which deEnes the yielding of the material. The material properties, Young\u2019s modulus (E) of 200GPa, Poisson\u2019s ratios (n) of 0.25, 0.35 and 0.45, yield strength (Y) ranges from 210MPa to 2520MPa (Malayalamurthi and Marappan,14 Kogut and Komvopoulos,15 Chatterjee and Sahoo16) and tangent modulus(Et) of 0E, 0.02E, 0.1E and 0.2E (Sahoo and Chatterjee,10 Kadin et al.17) are considered. The meshed model is shown in Figure 1. The extreme left side nodes of the asperity are given with axisymmetric boundary condition. The displacement is applied from top nodes of the asperity in incremental manner. The rigid flat plane is restricted to move in all directions. The meshed model is validated with analytical Hertz\u2019s elastic solution and found less than 1% deviation. Then, the mesh density is doubled until the contact force and contact area are differed by 1% between the iteration for elastic-perfectly plastic material case (Et\u00bc 0E), and the results are validated with JG, SM and MM elastic-perfectly plastic models"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001276_isgt-la.2019.8895013-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001276_isgt-la.2019.8895013-Figure1-1.png",
        "caption": "Fig. 1. Schematic diagram of interturn faults.",
        "texts": [
            " Thus, there are opportunities for investigations regarding the application of machine learning techniques to PMSG generators, which are used in wind turbines. One of the problems of such type of research is the lack of training data from real turbines since supervised learning techniques highly dependent on consistent data to the creation of accurate models. III. RESEARCH METHOD One of the most common types of internal faults in electric machines is the leakage current of the coils through connections, which were caused by faults in the insulation of the components. In Fig. 1 these connections have an electrical resistance Rscj, with j = a, b and c. The value of this resistance depends on the insulation wear, being Rscj = \u221e for no faulty and Rscj = 0 for total insulation disruption conditions. Besides, different fractions of coil turns nj can be short-circuited [17]. The modelling of the internal fault must consider the effects of a fault on the resistances and inductances of the faulty coils. In order, the stator coil resistance can be given by (1), [17]. = (1) where \u03c1 - resistivity of the coil material; l - coil wire length; The only parameter sensitive to interturn fault is l. Thus, if a fraction nj of the coil turns are short-circuited to each other, resistance equal to njRs matches to the faulty fraction of the coil and another resistance equal to (1-nj)Rs matches to the no faulty fraction. Self-inductance of solenoids like the coils shown in Fig. 1 can be computed through (2), [17]. = (2) where \u03bc - magnetic permeability of the coil core; N - number of coil turns; A2 - rectangular area closed by the coil wire. The parameters sensitive to interturn faults are N and l. Therefore, similarly to the resistance, an inductance equal to njLs matches to the faulty fraction of the coil and another inductance equal to (1-nj)Ls matches to the no faulty fraction. Since the mutual inductance between stator coils Ms is given by (3), this parameter suffers a similar effect than Ls [17]"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000985_j.msea.2019.05.017-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000985_j.msea.2019.05.017-Figure1-1.png",
        "caption": "Fig. 1. Schematic of the experimental geometry employed in this work with the laboratory and sample coordinate systems denoted with L and S respectively. Incoming X-rays travel in the k\u0302i direction and diffracted X-rays travel along k\u0302o. Measured diffracted intensity is parameterized by three angles \u03b82 , \u03b7, and \u03c9. Radiographs are measured on a scintillator placed close to the specimen that can be moved out of the path of the X-ray beam and diffracted X-rays are measured on a large area detector behind the specimen.",
        "texts": [
            " To complete the X-ray measurements at a fixed material state, the mechanical loading was paused and the applied load was reduced to 75% of the maximum load to prevent creep. At each load step, three 1mm tall volumes were probed along the gauge length. However, the differences in the initial texture and mechanical response of these volumes were found to be negligible (as expected from a uniform gauge) so results will focus on the center volume. A schematic of the experimental geometry is shown in Fig. 1. The laboratory and sample coordinate systems are denoted with the superscripts L and S respectively. The laboratory coordinate system is defined such that the incoming beam direction k\u0302i is equal to \u2212 ez L and the rotation/load axis of the specimen lies parallel to ey L. The energy of the incoming X-ray beam was 61.332 keV (wavelength =\u03bb 0.0202 nm) and was 2mm wide (along ex L) and 1mm tall (along ey L). The beam was made sufficiently wide to illuminate the entire cross section of the specimen as it rotated",
            " From the fit peak data, the intensity of each peak and the centroid \u03b82 values are calculated. The intensity data is used to calculate the probability of finding lattice planes of a given orientation for the orientation pole figure and the \u03b82 values are used to calculate strain. As the last step, measured scalar values for each peak in an azimuthal bin are then mapped to a sample direction on the unit sphere. The mapping from diffraction angles to a direction in the sample frame q\u0302S in the experimental geometry shown in Fig. 1 is [19]: = \u23a1 \u23a3 \u23a2 \u23a2 \u2212 \u23a4 \u23a6 \u23a5 \u23a5 \u23a1 \u23a3 \u23a2 \u23a2\u2212 \u2212 \u23a4 \u23a6 \u23a5 \u23a5 q d \u03bb \u03c9 \u03c9 \u03c9 \u03c9 \u03b8 \u03b7 \u03b8 \u03b7 \u03b8 [ \u02c6 ] cos( ) 0 sin( ) 0 1 0 sin( ) 0 cos( ) sin(2 )cos( ) sin(2 )sin( ) cos(2 ) 1 S hkl (1) where dhkl is the spacing of a set of lattice planes. The inversion of orientation pole figure data to calculate orientation distribution functions (ODFs) was performed using the ODFPF software package and a procedure detailed in Ref. [20]. The package uses finite elements to represent functions over pole figures and orientation space (specifically Rodrigues space)",
            " For a set of lattice planes with normals oriented along the direction q\u0302, their current lattice strain \u03b5hkl is defined as = \u2212 = \u2212\u03b5 d d d \u03b8 \u03b8 sin( ) sin( ) 1hkl hkl hkl hkl 00 0 (7) where 0 indicates an initial or unstrained value. From shifts in the Bragg angle, the strains of differently oriented sets of lattice planes are determined and mapped to the unit sphere in the same fashion as more traditional orientation pole figures. Importantly for interpretation, different subsets of crystals contribute to each point on the strain pole Fig. 1 a material is elastically or plastically anisotropic at the crystal scale, strain pole figures from different families of lattice planes hkl( ) and lattice planes of varying orientation within the same family will exhibit varying mechanical responses. More information about the relationship between orientation pole figures, strain pole figures, and the underlying distributions of these quantities in orientation space is detailed in Refs. [31,42,43]. [1] W.E. Frazier, Metal additive manufacturing: a review, J"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0003659_mssp.1998.0190-Figure8-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0003659_mssp.1998.0190-Figure8-1.png",
        "caption": "Figure 8. Test structure showing the six response measurement points and the six excitations.",
        "texts": [
            " (2) Select the excitations so that their extension lines do not meet each other. (3) Select the excitations so that their extension lines are located as far as possible from the centre of gravity. (4) Before identi\"cation of the moments of inertia, check the scalar triple-product term of the angular acceleration vectors, and, if it is small, then change excitation directions or points. 6. EXPERIMENTS All the results suggested in the preceding sections were tested with a three-dimensional arbitrarily shaped structure as shown in Fig. 8. The test structure was composed of three aluminium plates attached to \"ve rubber mounts. The loss factor of the rubber mounts was 0.08}0.1. As shown in Fig. 8, the six excitation conditions and the six response measurement points were chosen as candidates for excitation and response points. Then the 108 frequency response functions between the prechosen six response measurement and the six excitation points were measured with an impact gun and a triaxial accelerometer. The measured natural frequencies of the six rigid-body modes were 7.8, 11.9, 16.0, 20.0, 24.7 and 29.6 Hz, and the natural frequencies of the \"rst two elastic modes were 167.0 and 212.5 Hz",
            " Figure 9 shows an example of measured and curve-\"tted FRF. As shown in the \"gure, the direct estimation of the mass line from the measured FRF is not feasible, since the elastic modes considerably in#uence the mass line. After curve-\"tting all the measured FRFs, the 108 inertia restraints were obtained. When the hand-calculated inertia restraints are considered as true ones, the experimentally obtained inertia restraints have a 3}10% relative error (i.e. D error D/true value D), respectively. Using the test structure shown in Fig. 8, the suggested guides for the selection of response measurement points were tested. Three measurement points were chosen from six prede\"ned measurement point candidates on the test sturcture; Table 1 lists the positions of the three selected response measurement points for each case. Figure 10 shows a very close correlation between the two values, the condition number of [R p ] and the ratio of length of area. Among all of the possible measurement cases, &I (P a , P b , P c )', &II (P a , P d , P e )', and &III (P a , P d , P f )' have relatively small condition numbers; at the same time, those are the cases for which the three measuring points form almost a regular triangle",
            " The origin point accelerations were calculated using experimentally measured inertia restraints and compared with true values. Figure 11 compares the relative errors of accelerations at the origin point (i.e. E Mda o N E/E Ma o N E) versus measurement selection cases. Even though the error varies with excitations, the relative error is roughly proportional to the ratio of the length to the area. This means that the ratio of the length to the area can be a good guideline in selecting a set of response measurement points. Next, the six excitation candidates plotted in Fig. 8 on the same test structure were used to examine the relationship between the excitation conditions and the condition numbers of [R r ] and [R q ]. Three excitations were chosen from a set of six prede\"ned excitations. Table 2 lists the chosen excitations versus excitation case numbers. Figures 12 and 14 compare the condition numbers [R r ] and [R q ] with the magnitude to area and magnitude to volume ratios, and show that the physical quantities are roughly proportional to the condition numbers"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001217_j.mechmachtheory.2019.103633-Figure15-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001217_j.mechmachtheory.2019.103633-Figure15-1.png",
        "caption": "Fig. 15. Force analysis of crank-4BSL.",
        "texts": [
            " It can be seen from Fig. 14 that the maximum error between the simulation and theoretical results is less than 1.5%, which verifies the validity of the load-displacement relationship mode of rectangular spring leaves. Therefore, a ZSFH can be constructed with a rectangular leaf spring string. According to the stiffness characteristic of 4BSL in Section 3.1 , the K 4BSL of 4BSL has the following equation: F \u03b3 = K 4 BSL x \u03b3 (8) K 4BSL is the stiffness of the 4BSL. The force analysis of the crank-4BSL is as shown in Fig. 15 . The initial angle \u03b2 is the angle between AB \u03b3 and AC when the spring is in the natural state (i.e. the spring is not deformed). M \u03b3 is the moment of the force on the crank, and the crank rotates from AB \u03b2 to AB \u03b3 under the function of M \u03b3 . \u03b3 is the crank angle between the crank and the x-axis during rotation of the crank, r is the crank length, and l is the base length. It is defined that \u03be = r/l (0 < \u03be < 1). From the geometric relationship, it can be known that: { \u2223\u2223B \u03b2C \u2223\u2223 = l \u221a \u03be 2 + 1 \u2212 2 \u03becos \u03b2\u2223\u2223B \u03b3 C \u2223\u2223 = l \u221a \u03be 2 + 1 \u2212 2 \u03becos \u03b3 (9) The torque equilibrium equation to crank AB \u03b3 is: M \u03b3 = F \u03b3 d \u03b3 = \u03be l sin \u03b3\u221a \u03be 2 + 1 \u2212 2 \u03becos \u03b3 F \u03b3 (10) Where d \u03b3 is the force arm of F \u03b3 to point A, F \u03b3 is the equivalent restoring the force of the 4BSL",
            " The relationship of torque-crank angle (include dimension) of the IORFH is [30] : M IORFH = E IORFH _ B I IORFH _ B L IORFH _ B [ 168 \u03b82 d ( 9 \u03bb2 \u2212 9 \u03bb + 1 )2 6300 + \u03b82 d ( 9 \u03bb2 \u2212 9 \u03bb + 11 ) + 12 ( 3 \u03bb2 \u2212 3 \u03bb + 1 )] \u03b8 (16) Where, E IORFH _B is the elastic modulus of the material, L IORFH _B is the length of the beam flexure of IORFH, I IORFH _B is the moment of inertia of the section, \u03bb is the ratio of the distance from the intersection to the short end and the length L IORFH _B of the beam flexure of IORFH. When, \u03bb = 0.1273, the relationship of torque ( M IORFH )-rotation angle ( \u03b8 ) is linear. In Fig. 15 , the crank angle \u03b3 of the negative stiffness mechanism and the rotation angle \u03b8 of the IORFH are exactly the same angle, that is: \u03b3 = \u03b8 (17) By putting the Eqs. (10) , (15) , (16) and (17) together, and putting \u03bb = 0.1273 into the Eq. (15) , we can get the relationship of the restoring force F \u03b3 of the negative stiffness spring (4BSL) and the rotation angle \u03b3 of IORFH, that is: F \u03b3 = 8 E IORFH _ B I IORFH _ B \u03b3 n L IORFH _ B \u03be l sin \u03b3 \u221a \u03be 2 + 1 \u2212 2 \u03becos \u03b3 (18) According to the Eq. (17) and the geometric relationship of the spring-crank mechanism in the process of deformation, we can get the relationship between \u03b3 and x \u03b3 , that is: \u03b3 = arccos { 1 2 \u03be [ \u03be 2 + 1 \u2212 (\u221a \u03be 2 + 1 \u2212 2 \u03becos \u03b2 \u2212 x \u03b3 l )2 ] } (19) Therefore, the relationship of restoring force F \u03b3 and rotation angle \u03b3 based on the spring-crank mechanism is expressed by Eqs"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000234_ecce.2019.8913191-Figure9-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000234_ecce.2019.8913191-Figure9-1.png",
        "caption": "Fig. 9. Flux distributions of comparative VFMM models when applying positive magnetizing current pulse of 600 Arms.",
        "texts": [
            " 8(a), VPMs of Model-A are fully demagnetized at around -600 Arms. In contrast, VPMs of Model-A are almost not remagnetized despite applying +600 Arms. This is because VPMs are affected by diamagnetic field caused by CPMs in V-type CPMs arrangement. As a result, negative and positive magnetizing currents are unbalance in Model-A. In addition, most positive d-axis fluxes through CPMs when applying positive current pulse of +600 Arms, and therefore, magnetic flux density of VPMs are very low as indicated in Fig. 9(a). In addition, it is obvious that Model-A doesn\u2019t achieve even magnetic reversal. In Model-B having delta-type CPMs arrangement, operating point of VPMs when applying -100 Arms is stopped in front of the knee point unlike Model-A. This means that durability against unintentional demagnetization by load current of Model-B is improved, compared with Model-A. Additional CPMs of Model-B contribute to increasing durability against demagnetization of VPMs because magnetic fluxes of V-type CPMs as the diamagnetic field are gathered to the additional CPMs. Compared with Model-A, it is found that magnetization characteristic of Model-B when applying positive current pulse of +600 Arms is improved because magnetic reversal of VPMs is achieved as shown in Fig. 8(b) and Fig. 9(b). From the above, it is obvious that deltatype CPMs arrangement is effective in both increasing durability against unintentional demagnetization and improving re-magnetization characteristic. Nevertheless, VPMs of Model-B are not also fully magnetized by current pulse of +600 Arms, and therefore, more additional countermeasures are needed. Accordingly, Model-C employs the extended flux barriers in order to improve drawbacks of Model-A and B in terms of re-magnetization characteristics. It has been found that the residual magnetic flux density of VPMs is approximately 1.9 T when maximum current pulse of +600 Arms is applied as indicated in Fig. 8(c). This means that the magnetization state of VPMs can be changed to over +90% within three times the rated current. Fig. 9(c) shows the magnetic flux distribution of Model-C, and positive d-axis fluxes are obviously concentrated to VPMs owing to the extended flux barriers. In addition, durability against unintentional demagnetization is similar to Model-B. Consequently, Model-C can improve asymmetric positive and negative current pulses to regulate the magnetization state of VPMs. Summarizing the above, delta-type CPMs arrangement and the extended flux barriers are very effective methods against improvement of the magnetization characteristic of the VFMM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002789_0309524x21999841-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002789_0309524x21999841-Figure1-1.png",
        "caption": "Figure 1. Vibration monitoring for wind turbine drive train.",
        "texts": [
            " Vibration monitoring for wind turbine drivetrain For horizontal axis wind turbines, stochastic winds are absorbed by blades and wind energy is converted into mechanical energy of rotor hub with low rotational speed, then the low rotational speed is speeded up through a gearbox to drive the generator in high speed. The blades and rotor hub are supported by main bearing. The rotor hub, main shaft, gearbox, and generator form the drive train. Several accelerometers are attached on different positions to monitor the health status of the drive train, shown as in Figure 1. For the accelerometer 1 in Figure 1, it not only can monitor the status of main bearing, but is beneficial to analyze the balance of the rotor system, because it is the closest one near the rotor hub. Vibration model of rotor imbalance with faulty blade Vibration signal is the reflection of the health status of wind turbine blades and drivetrain. Through processing the vibration signal collected from the drivetrain, a quantity of faults, for example, rotor imbalance, gears or bearings crack, etc., can be detected. In view of this, the fault characteristics hidden in vibration signal need to be analyzed first",
            " The tested wind turbine is a fixed-pitch one, the rated power of which is 750 kW. The wind turbine were tested four times in all. Among these, blade crack was found during the second test. The first test was implemented 1 year before the second test. The third test was immediately carried after repairing the crack of the faulty blade. The forth test happened after the third one. The rotational frequency of the rotor hub was 0.362 Hz during the test. The test was offline, one accelerometer by one on the surface of the wind turbine drive train, shown as in Figure 1. The sampling frequency of accelerometer 1, 2, and 3 in Figure 1 is 5120 Hz due to the relatively low rotational speed of the monitored parts, and the sampling frequency of the rest ones is 25,600 Hz. Figure 4 shows the vibration signals from accelerometer 1 during the four tests. Figure 4(a) to (c), and d correspond with the four vibration tests sequentially. All the vibration amplitudes in Figure 4 are in the range of 64 m/s2, where the amplitude in the second test with blade crack is lower than the other tests. This brings an opposite result to the original viewpoint that faults will make vibration larger"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000522_j.mechmachtheory.2016.04.002-Figure7-1.png",
        "caption": "Fig. 7. Characteristics of the 6R mechanism.",
        "texts": [
            " The above results have been verified using several mechanism models built using 3D printing. Fig. 6 shows the CAD model and 3D-printed prototype of 6R Mechanism I (Fig. 4c). It is noted that joints 1 and 6 in this prototype are prevented from full-cycle rotation due to interference between links 2 and 4 as well as links 1 and 5. Let K1, K2 and K3 denote the intersections of joint axes of joints 1 and 6, joints 2 and 5, and joints 3 and 4. P1, P2 and P3 represent the plane defined by the axes of joints 1 and 6, joints 2 and 5, and joints 3 and 4 respectively (Fig. 7). From Ref. [36], we obtain that planes P1,P2 and P3 and plane K1K2K3 have a common point, K, at any configuration of the 6R mechanism during motion. It is noted that at configurations A, C (Fig. 4), D and F (Fig. 5) of Mechanism I, point K2 is at infinity. A 6R mechanism that has three pairs of R joints with intersecting joint axes has been proposed using a geometric construction approach. Kinematic analysis of the mechanism has been presented. The analysis has shown that the 6R usually has two solutions to the kinematic analysis for a given input"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002770_j.jmapro.2021.02.015-Figure6-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002770_j.jmapro.2021.02.015-Figure6-1.png",
        "caption": "Fig. 6. The results of the experiments: (a) 90\u25e6 rotate scanning strategy; (b) 67\u25e6 rotate scanning strategy; (c) \u201cskin-core\u201d scanning strategy with 90\u25e6rotate \u201ccore\u201d scanning; (d) \u201cskin-core\u201d scanning strategy with 67\u25e6rotate \u201ccore\u201d scanning.",
        "texts": [
            " Parameters Scanning speed Laser power Hatch space Layer thickness Value 1000 mm/s 300 W 0.1 mm 40 \u03bcm Table 4 The designed dimensions of the samples. Features External diameter Internal diameter Wall thickness Height Dimension 40 mm 36 mm 2 mm 12 mm L. Zhang et al. Journal of Manufacturing Processes 64 (2021) 907\u2013915 Table 3 were used. The designed dimensions of the samples were shown in Table 4. Finally, four circle structure samples with different scanning strategies were fabricated, as shown in Fig. 6. Fig. 7 shows the research flow chart. In order to reduce the thermally stress and avoid the deformation, the SLMed samples were heat treated after the forming process. Firstly, the samples were put into a vacuum furnace at 500 \u2103 with the heating rate of 8.3 \u2103/min. And then the temperature was maintained 2 h and furnace cooling at last. After the heat treatment, the samples were cut from the substrate by a wire electrical-discharge machine and cleaned by the sandblasting [30] and ultrasonic vibration before the measurement to reduce the adhered powder on the surface, which would increase the measured results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000564_j.eurpolymj.2016.05.016-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000564_j.eurpolymj.2016.05.016-Figure11-1.png",
        "caption": "Fig. 11. (a) SECM approach curves for conductive (black) and insulating (red) substrates. (b) Microelectrode in bulk solution indicating the hemispherical diffusion profile; microelectrode in close vicinity of an insulating surface (negative feedback effect) and in close vicinity of a conductive surface (positive feedback effect). (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)",
        "texts": [
            " Reducing the size of an electrode in at least one dimension regardless of a disc, ring, band or conical electrode shape frommacroscopic to microscopic dimensions leads to a significant change in the electrochemical behavior resulting from the fact that the diffusion layer is significantly larger than this particular dimension of the microelectrode. The diffusion contributions towards the edge of the electroactive area of a disc microelectrode results in a hemispherical diffusion profile in contrast to the planar unidirectional diffusion towards a macroscopic disc electrode (see Fig. 11b (left) and Fig. 2b, respectively), which facilitates the enhanced mass transport towards the active micro-sized electrode surface. Mathematically, this is expressed by the modified Cottrell equation comprising a time-dependent transient term (Eq. (1)), with a steady state term (Eq. (2)), in contrast for the time dependent current at a macroscopic electrode. Please Polym i\u00f0t\u00de \u00bc nFAc ffiffiffiffi D p ffiffiffiffiffi pt p \u00f01\u00de i\u00f0t\u00de \u00bc nFAc ffiffiffiffi D p ffiffiffiffiffi pt p \u00fe 4nFDcro \u00f02\u00de where i(t) is the time dependent Faraday current, n is the number of transferred electrons, F is the Faraday constant, D is the diffusion coefficient, A is the electrode area, c is the concentration of the electroactive species, t is the time, and ro is the radius of the microelectrode",
            "016 samples with considerable changes in height and high-aspect-ratio features; a recent review of positioning modes can be found elsewhere [72]. In traditional SECM experiments, the probe is positioned by adding a redox species exhibiting fast electron transfer characteristic in the electrolyte solution. The faradaic current at the probe, which is biased to oxidize or reduce the redox species is recorded while the probe is approached in z-direction towards the sample surface. Such socalled approach curves in dependence of the conductivity of the sample are shown in Fig. 11a. If the probe is in bulk solution, i.e., far away from the sample surface and only diffusion is considered, the recorded steady state current of a disc-shaped electrode can be expressed by Eq. (9) (see Section 2.2.1. [15]). The approach is stopped once the concentration profile \u2013 and hence the Faraday current recorded at the probe \u2013 is apparently influenced by the presence of the sample. Based on numerical simulations of reaction-diffusion equations with appropriate boundary conditions, the SECM response current can be quantified in dependence of the involved redox species, the electroactive area of the probe, and the thickness of the electrode \u2013 usually glass \u2013 sealing (RG value: rg/rT ratio). Mapping and investigating biologically relevant redox active macromolecules, e.g., investigating DNA-modified surfaces and mapping enzyme activity is usually obtained in feedback (FB) [73] or generation-collection (GC) [74] mode. For mapping local oxygen consumption, frequently the redox competition mode is used [75]. The feedback mode of SECM is shown in Fig. 11b. A redox species, e.g., its reduced form Red is added to solution. Biasing the probe sufficiently to achieve a steady state current, the Faraday current is increased if the probe is positioned in close proximity (i.e., several electrode radii) above a conductive sample Fig. 11b. The reduced species, Red is oxidized at the probe and Ox diffuses within the gap to the conductive surface where it is regenerated to its initial redox state, Red. Hence, locally an increased concentration of the reduced form of the redox species is obtained resulting in a positive feedback current, and the probe current increases with decreasing probe-surface distance. This can be expressed by following analytical expression for an infinitive conductive substrate and a RG value of 10 [76]: Please Polym icondT \u00f0K\u00de \u00bc iT iT;1 \u00bc 0:72627\u00fe 0:76651 K \u00fe 0:26015e 1:1:41332=K \u00f023\u00de where K is the normalized distance, (K = d/ro)"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002903_j.engfailanal.2021.105453-Figure21-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002903_j.engfailanal.2021.105453-Figure21-1.png",
        "caption": "Fig. 21. Crack location.",
        "texts": [
            " Based on the above analysis, it can be clearly understood that the failure process of the rubber joint is as follows: first of all, due to the failure of the rubber joint during the installation process is exposed to corrosive detergents, resulting in the internal filler to precipitate out and make the rubber swelling. The precipitation of filler and swelling of rubber weaken the resistance to aging and external forces. Therefore, when it is subjected to cyclic external forces, it will appear a certain degree of fatigue (When the subway is running, it is inevitable that the road surface will be unsmooth. Therefore, the rubber joint will inevitably vibrate up and down with the movement of the subway, which lead the cracks occur in the position shown in Fig. 21.). Secondly, because of the presence of too large C. Su et al. Engineering Failure Analysis 126 (2021) 105453 filler particles and aggregation of small particle fillers in rubber at the crack, there is stress concentration at both the large particle and small particle agglomeration, which leads to the accelerated development of fatigue crack at the stress concentration. In addition, rubber itself has the problem of excessive crosslinking density. Excessive crosslinking density leads to excessive hardness and insufficient elasticity of rubber, which further aggravates the speed of fatigue crack propagation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000578_978-4-431-55879-8_3-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000578_978-4-431-55879-8_3-Figure4-1.png",
        "caption": "Fig. 4 The robot i explores U2 \u2286 V k j to check if k can reach it. The exploration path is shown with dark dashed line ending with an arrow. a U2 is unreachable to the robot k, hence it is added to Si . b Though U2 is reachable to the robot k, it has to pass through U1 to reach it. In other words, U2 is not adjacent to V k (V k \u2229 U2 = \u2205). Thus U2 \u2208 Si . c The robot i reaches a point on A f k j while exploring U2 and hence U2 /\u2208 Si (as U2 \u2208 Sk )",
        "texts": [
            " Let u jk(m) \u2208 Ab jk contains Pi jk (Such u jk(m) exists as Pi jk is assumed to be part of U2 adjacent to U1). 1a. If u jk(m) \u2229 A f k j = \u2205, as illustrated in Fig. 3c, a, then k can reach U2, and hence i will not cover it. 1b. Otherwise, as illustrated in Fig. 3c, b, k can not reachU2 and i should cover it. The robot i can check if Pi jk \u2208 U1 \u2229 U2, and ifU1 \u2229 U2 is a single connected piece, while physically exploring the boundary of U1. 2. Consider a scenario, U1 \u2229 V k j is a not a single line segment or Pi jk /\u2208 U1, as illustrated in Fig. 4. In such a scenario, robot i will not be able to decide if U2 needs to be added to Si or not only based on available information. The patch U2 is added to Si , only if, while physically exploring the boundary of U2, the robot i reaches a portion of A f k j . Remark 1 Note that the robot i physically explores the boundary of a patch W which is adjacent to U \u2208 Si , only when the information about free and blocked regions of Voronoi cell boundaries (Ai j ) is not sufficient to make a decision as to if W needs to be added to Si . Such an exploration is local to the robot i and it does not affect the decisions of other robots. This can be observed from the illustrations in Fig. 4. The patch U2 is added to Si (Fig. 4a and b) when i concludes that U2 /\u2208 Sk , and is not added to Si (Fig. 4c) when U2 \u2208 Sk . This ensures that the patch U2 is covered exactly by one robot. Further, it can be noted the scenarios discussed above are exhaustive. Scenario iii. Patches in V l j , j \u2208 N (i), l \u2208 N ( j), l /\u2208 N (i) If V i j \u2283 U1 \u2208 Si and U2 \u2282 V l j , s.t. U2 is adjacent to U1, then robot i has to make a decision on adding U2 to Si . (1) If V l j \\V j0 j is not accessible to robot l (based on A f l j and Ab jl , discussed in scenario(i)), then U2 is added to Si . Such scenarios are illustrated in Fig"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001350_0954407019890481-Figure11-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001350_0954407019890481-Figure11-1.png",
        "caption": "Figure 11. Submodel of cylinder head.",
        "texts": [
            " The calculated results showed a good agreement with the measurement. Their trends are the same. With increasing distance to the cylinder center, its temperature decreases. The measured results are slightly higher in magnitude than the calculated results. The error is less than 10%, except for measuring points 4 and 7. The points near the center have relatively larger errors. It implies that these points are more critical and caused more challenges in simulations. The thermal and displacement boundary conditions of the submodel, as shown in Figure 11, were mapped from the global model results. Figure 12 shows the transient thermal load conditions on the submodel, where P0 represents for assembly step, P5 represents for hot assembly step at rated speed, and P9 represents for cold assembly step at idle speed. The submodel temperature and displacement at P0, P5, and P9 load conditions were measured by the static stress\u2013strain calculation of the global model, and the other thermal load and boundary conditions were calculated by the following equations T i\u00f0 \u00de=T5 T5 T9\u00f0 \u00de3K i\u00f0 \u00de \u00f03\u00de U i\u00f0 \u00de=U5 U5 U9\u00f0 \u00de3K i\u00f0 \u00de \u00f04\u00de where T(i) is the temperature of the submodel, U(i) is the displacement of the outer surface of the submodel, and K(i) is the temperature and displacement coefficient from test results"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001546_s12206-020-0209-1-Figure7-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001546_s12206-020-0209-1-Figure7-1.png",
        "caption": "Fig. 7. Engine power mode selection flow chart.",
        "texts": [
            " In this study, the research group cooperated with a company to develop a 5-ton wheel loader prototype as a test case. After negotiation with the engine manufacturer, the engine provided three power curves, namely light load, middle load, and overload, as shown in Fig. 6. The engine mode is manually selected by the driver using the engine power mode knob in the cockpit. Before the driver's shovel maneuver, the required engine power is determined based on the driver's experience, and the engine power mode knob is adjusted to be in the proper gear position. Fig. 7 shows the flow chart. Different gear positions correspond to distinct resistance values to change the current of the input controller and select different engine power curves. Fig. 7 demonstrates that the engine power mode knob has three gear positions of different resistance. After selecting gear position, the current input to the engine controller changes and the corresponding engine power curve is selected by the controller. The above energy-saving operation is affected by the driver's subjective awareness and operating habits. Thus, the utilization rate is not high. However, it is required to improve properly the engine energy-saving control of the wheel loader in that the engine controller can automatically select the appropriate working mode without the impact of human operator"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002218_j.mechmachtheory.2019.103771-Figure4-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002218_j.mechmachtheory.2019.103771-Figure4-1.png",
        "caption": "Fig. 4. Hoeckens four-bar linkage.",
        "texts": [
            " In Section 3 , the rigid-body dynamic and the elastodynamic models are constructed, respectively. Section 4 presents the solutions to the elastodynamic equations and the elastodynamic response curves of this novel forging manipulator. Conclusions are provided in Section 5 . In this section, a novel parallel-link mechanism for forging manipulators is proposed, and its main-motion mechanism is shown in Fig. 3 . The horizontal buffering mechanism is composed of an equivalent mechanism of the Hoeckens straight-line mechanism (see Fig. 4 ) when the buffering cylinder q 3 is locked, and the ratio of the length of each linkage satisfies l PN : l PH : l N G : l MG : l GH = 1:2:2.5:2.5:2.5 [24 , 25] . The output point ( M ) of the Hoeckens straight-line mechanism is directly connected to the middle part of the gripper carrier ( EF ), and the buffering cylinder q 3 is installed into the linkage HG of the straight-line mechanism to form a horizontal buffering mechanism. The differences between the main-motion mechanisms of novel parallel-link manipulator and typical ones are the installation location and the architecture of the horizontal buffering mechanism. For the typical manipulators, the horizontal buffering mechanism is composed of only a buffering cylinder whose fore-end is connected to the middle part of link CE , and the rear-end to the truck frame. The Hoeckens straight-line mechanism is a kind of mechanism which can transform the rotating motion of the input into an approximately straight-line motion of the output. When the crank PN rotates 360 \u00b0 anticlockwise around point P from the initial position shown in Fig. 4 , the trajectory of the output point M is shown in Fig. 5 . It can be found that when PN rotates in the range of \u221290\u201390 \u00b0, the trajectory of point M is nearly a straight line (the maximum deviation in y -direction is less than 1 mm when l 2 = 100 mm). Thus, the motion of the output point M can be regarded as a straight-line motion when the crank angle meets the specific conditions given above. For this novel forging manipulator mechanism, when the gripper carrier is driven by a hydraulic cylinder for lifting motions within a particular scope, the coupled displacement of the gripper carrier in the horizontal direction is almost zero, thus realizing the decoupling between the lifting and the horizontal translation"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000683_tmag.2016.2601886-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000683_tmag.2016.2601886-Figure12-1.png",
        "caption": "Fig. 12. Equipotential distribution at aligned position with injecting rated qaxis currents, PMs are replaced by air.",
        "texts": [
            " 11 shows the dq-axis inductances of SS-PMM and two proposed PS-PMMs with the same 12/10 stator/rotor pole number combination against different current angles with the rated currents as given in Table I and Table II. Similar to SSPMM, the d-axis inductances are also quite close to q-axis inductances in both PS-PMM-I and PS-PMM-II. Consequently, the saliency ratios are all close to 1. Therefore, the potential reluctance torques can be negligible in both PS-PMM-I and PS-PMM-II. Moreover, both PS-PMM-I and PS-PPM-II exhibit higher dq-axis inductances than SS-PMM due to the shorter equivalent air-gap length in main magnetic flux path as shown in Fig. 12. E. Self- and Mutual-Inductances Fig. 13 shows the variation waveforms of self- and mutualinductances against with rotor positon for all machines with 12/10 stator/rotor pole number combination. pa = 10W is corresponding to the copper loss produced by the dc current, which is injected into the phase A winding. Obviously, both PS-PMM-I and PS-PMM-II exhibit higher self- and mutualinductances, which are both 2.7 and 3 times of those for SSPMM, respectively. It is also due to the shorter equivalent airgap length in main magnetic flux path of PS-PMMs when compared with SS-PMM"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001357_j.oceaneng.2019.106812-Figure12-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001357_j.oceaneng.2019.106812-Figure12-1.png",
        "caption": "Fig. 12. Simplified Cylinder Model with bead-on-plate welding.",
        "texts": [
            " In order to understand the mechanism of cross section change behavior of examined cylindrical leg structure after welding fabrication, a simplified cylinder model with bead-on-plate welding on outer surface was investigated in advance, then due to symmetrical feature and to reduce the computation consumption, quarter FE model of rack and cylinder welded structure was created to predict the welding deformation and compare with experimental measurement. An identical dimension according to cylindrical leg structure without rack component was considered as shown in Fig. 12, and the total number of points and elements are 37,824 and 25,767, respectively. Multi-pass bead-on-plate welding lines were performed on the outer surface, and the length of welding line was assumed as 500 mm. Welding conditions as summarized in Table 1 and material properties as shown in Fig. 13 are all identical comparing with experiment. The inter pass temperature is considered as about 200 \ufffdC for later FE computation. Fig. 14 shows the contour plotting of highest temperature distribution as well as welded zone, it also can be seen that welding arc has a little bit shallower penetration in the thickness direction comparing to the thickness of cylinder",
            " It also can be seen that the cross section shape has welding distortion with shrinkage in upright direction and expansion in horizontal direction. From the above computational results and explanation, the mechanism of welding distortion generated during rack and cylinder joining on the cross sectional shape of cylindrical leg structure can be clearly understood. Meanwhile, in order to demonstrate the efficiency of proposed methods during TEP FE computation, computational times with the FE model as shown in Fig. 12 were summarized and compared in Table 2. In-house Fortran code was compiled with Intel Parallel Studio XE 2011 based on the platform of Ubuntu 14.04 OS and Dell PowerEdge T420 Server. It can be concluded that computing time during the thermal analysis can significantly be reduced to half using Intel Math Kernal Library (Intel MKL) or OpenMP Parallel Computation with 10 threads, and to quarter using both Intel MKL and OpenMP Parallel Computation with 10 threads. With much more TEP FE computations, computational efficiency was compared not only for thermal analysis but also for mechanical analysis as summarized in Table 2"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0002161_9783527813872-Figure7.19-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0002161_9783527813872-Figure7.19-1.png",
        "caption": "Figure 7.19 (a) Scheme of a zone-casting apparatus and vertical phase separation mechanism and (b) TEM image of phase-separated PTCDI-C5 and polycarbonate film. Source: Dobruchowska et al. 2014 [79]. Reproduced with permission of Elsevier.",
        "texts": [
            " During the deposition, a directional concentration gradient occurs in the meniscus, which subsequently induces a uniaxial crystallization of the organic semiconductor in the area of increased concentration. Dobruchowska et al. applied zone-casting to fabricate blend films of n-type small-molecule bisethylpropyl-perylene-bis(dicarboximide) (PTCDI-C5 and insulating polycarbonate (PC)) (Figure 7.16) [79]. The films exhibited a high degree of vertical phase separation with PTCDI-C5 as the upper layer (Figure 7.19). The PTCDI-C5 molecules assembled into elongated, uniaxial ribbons with a high crystalline order. The phase separation was a result of two main factors. Firstly, the PC solidified before PTCDI-C5 when the concentration reached a critical level in the meniscus. Furthermore, the small mixing entropy of both compounds resulted in expelling the PTCDI-C5 molecules to the top of the film. The three described types of interactions (solute\u2013substrate, solute\u2013solvent, and solute\u2013solute) are the main driving forces that lead to the vertical phase separation between the insulator and semiconductor"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0000073_icra.2019.8793549-Figure1-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0000073_icra.2019.8793549-Figure1-1.png",
        "caption": "Fig. 1. A prototype of the redundantly actuated omnicopter with eight propellers capable of producing forces/torques in all directions. The system is based on the design originally introduced in [11].",
        "texts": [
            " Other interesting implementations of fully actuated UAVs can be found in [9] and [10], but these designs also suffer from an unequal distribution of actuation within the force-torque space. A new UAV with even distribution of actuation is presented in [11]. It employs eight propellers evenly distributed about arms that span from the center to the vertices of a cube frame. The vehicle can follow an arbitrary position and orientation trajectory. The work presented here focuses on a new inverse actuator model and a new strategy for control allocation in the redundantly actuated UAV shown in Fig. 1, which is based on the design proposed in [11]. The work in [11] ignores aerodynamic interactions among the propellers in the control of the drone even though their effects on the vehicle performance can be significant. A major contribution of this paper is a new inverse actuation model that accounts for airflow interactions in relating the the propellers thrust forces to their motor commands. The inverse model is identified and validated experimentally and is used for real-time control of the vehicle"
        ],
        "surrounding_texts": []
    },
    {
        "image_filename": "designv11_14_0001237_j.mechmachtheory.2019.103652-Figure25-1.png",
        "original_path": "designv11-14/openalex_figure/designv11_14_0001237_j.mechmachtheory.2019.103652-Figure25-1.png",
        "caption": "Fig. 25. Test bench: (1) is the Wilcoxon 732A accelerometer; (2) is the valve; (3) is the cylinder head; (4) is the camshaft; (5) is the transmission gear.",
        "texts": [
            " The piston is not present in the test bench and an accelerometer Wilcoxon 732A is mounted on the end of a single valve to measure the motion. It is characterized by a measurement range of frequency from 0.5 Hz to 25 kHz. The accelerometer has a 10\u201332 female threaded connection, thus the connection to the valve is realized with a 10\u201332 to flat plug adapter bonded to a flat flanged connection, in turn, bonded to the valve with Loctite EA 9628H Aero cured at 121\u00b0C for 90 min in a climatic chamber. The piston of the engine is unmounted and the power is generated by an external electric motor, thus the proposed test bench (Fig. 25) is able to simulate only realistic geometric conditions. As in the real application, the valve is spring loaded to maintain the contact with the cam and it is dimensioned to avoid crossover shock. In the real application, a 0.08\u00f70.14 mm clearance is applied at the environmental temperature, to consider the natural expansion of the device (and especially of the valve) at the working temperature. In the considered test bench, this clearance is removed adjusting the screw valve, since the whole system works at the environmental temperature (neglecting heat transferred with friction)"
        ],
        "surrounding_texts": []
    }
]